- Email: [email protected]
Carbon dioxide degassing in fresh and saline water. II: Degassing performance of an air-lift
Aquacultural Engineering 43 (2010) 120–127
Contents lists available at ScienceDirect
Aquacultural Engineering journal homepage: www.elsevier.com/locate/aqua-online
Carbon dioxide degassing in fresh and saline water. II: Degassing performance of an air-lift Damian Moran ∗ DTU Aqua, Technical University of Denmark, Charlottenlund, Denmark
a r t i c l e
i n f o
Article history: Received 8 June 2010 Accepted 16 September 2010 Keywords: CO2 stripping Degassing Recirculating Reuse Salinity
a b s t r a c t A study was undertaken to measure the efficiency with which carbon dioxide was stripped from freshwater (0‰) and saline water (35‰ NaCl) passing through an air-lift at 15 ◦ C. The air-lift was constructed of 50 mm (OD) PVC pipe submerged 95 cm in a tank, had an adjustable air injection rate, and could be adjusted to three lifting heights: 11, 16 and 25 cm. The gas to liquid ratio (G:L) was high (∼1.9–2.0) at low water discharge rates (Qw ) and represented the initial input energy required to raise the water up the vertical riser section to the discharge pipe. The air-lift increased in pumping efficiency rapidly thereafter, to a G:L minima of 0.3–0.6 at 60–70 L min−1 . After this point the G:L ratio increased with Qw , representing decreasing air-lift pumping efficiency. The CO2 concentration of the influent and effluent water was measured using submersible infrared CO2 probes over a range of influent CO2 concentrations. The CO2 mass transfer coefficient [(kL a)20 ] ranged from 0.025 to 0.468. Increasing lift height increased mass transfer, which was attributed to both the increased G:L ratio and the contact time inside the air-lift. The relative effect of lift height and pumping rate on mass transfer was such that a 5 cm increase in lift height was approximately equal to a G:L increase of 0.5. The CO2 stripping efficiency was effectively the same between salinities, and the influent CO2 concentration only had a modest effect on CO2 stripping efficiency. At an influent concentration of 40 mg L−1 the CO2 stripping efficiency was 1–3% higher than at an influent of 10 mg L−1 . The relatively minor effects of salinity and influent CO2 concentration on stripping efficiency contrasted with a companion study investigating the stripping efficiency of a cascade column. The difference was attributed to the low-to-moderate mass transfer efficiencies of the air-lift. A general equation was derived for the airlift that allows one to calculate the mass transfer coefficient for a given lift height, Qw , or G:L ratio. The mass transfer coefficient can then be used to calculate the CO2 stripping efficiency for any water type (i.e. temperature, alkalinity, salinity and influent CO2 concentration). © 2010 Elsevier B.V. All rights reserved.
1. Introduction To design a system that meets the CO2 water quality requirements for a species, there needs to be knowledge about the biological tolerance of stock to CO2 , and the influx and efflux of inorganic carbon. This allows for a mass-balance analysis of the dynamics of inorganic carbon in the system to be carried out with a target operating CO2 concentration in mind. The aim of the current study was to provide data on the efflux of CO2 from an air-lift, a device used in aquaculture to pump water, ingas O2 and degas CO2 . Reinemann and Timmons (1989) and Timmons et al. (2002) provide an introduction to the theory and operation of air-lifts, but essentially an air-lift is a pipe submerged vertically into a water column with a 90◦ bend below, at or above the water surface. When
∗ Present address: Department of Biology, Lund University, Sölvegatan 35, S-22362 Lund, Sweden. Tel.: +46 46 22 27 785; fax: +46 46 22 24 425. E-mail address: [email protected]. 0144-8609/$ – see front matter © 2010 Elsevier B.V. All rights reserved. doi:10.1016/j.aquaeng.2010.09.001
air is injected at the bottom of pipe it rises and draws water with it, eventually discharging after rounding the 90◦ bend. To the author’s knowledge there is only one published study evaluating the CO2 degassing performance of an air-lift. Loyless and Malone (1998) measured the rate at which CO2 was removed from a tank of freshwater with an air-lift operating inside. While the CO2 removal rate was specific to the air-lift/tank configuration tested and could not be converted to a generalised measure of degassing efficiency (e.g. mass transfer or CO2 stripping efficiency), the study did show how lift height and air injection rate affected relative degassing performance. The current study adds to the work of Loyless and Malone (1998) using an almost identical air-lift to directly measure CO2 degassing performance, thereby allowing researchers to easily combine the lifting and pumping efficiency data of Loyless and Malone (1998) with the CO2 degassing efficiency data from this paper. A novel aspect of the current paper is the comparison of CO2 degassing efficiency in both fresh (0‰) and saline water (35‰ NaCl). Most other studies published to date on CO2 degassing have been solely concerned with freshwater (Grace
D. Moran / Aquacultural Engineering 43 (2010) 120–127
121
2. Materials and methods Nomenclature Alkt CO∗2 eq CO2(air)
CO2(in) non-eq CO2(eff ) eq
CO2(eff ) Ct Ct(in) Ct(eff) G:L kL a pCO2 Qg Qw t
total alkalinity sum of the unionized aqueous species H2 CO3 (aq) and CO2 (aq) CO∗2 concentration of water in equilibrium with air entering the air-lift CO∗2 concentration of influent water CO∗2 concentration of effluent water prior to equilibration reactions CO∗2 concentration of effluent water after equilibration reactions total inorganic carbon Ct concentration of influent water Ct concentration of effluent water gas:liquid ratio mass transfer coefficient (s−1 residence time) partial pressure of CO2 gas injection rate (L min−1 ) water flow rate (L min−1 ) time (s)
and Piedrahita, 1993, 1994; La Motta, 1995; Summerfelt et al., 2000, 2003; Watten et al., 2004).
1.1. Carbonate chemistry during CO2 degassing Grace and Piedrahita (1994) provide an extensive introduction to the gas transfer and chemistry processes that occur during CO2 degassing, and a companion paper describes how salinity influences CO2 degassing (Moran, 2010). A summary of the CO2 degassing process is as follows. As a body of water passes through a stripping unit, CO2 diffuses from the water into the air along the concentration gradient, a process known as mass transfer. The important difference between CO2 and other gases is that when CO2 dissolves in water, it undergoes chemical reactions, forming carbonates and H+ . The dissolution of CO2 into a number of aqueous species (CO2 (aq), H2 CO3 , HCO3 − and CO3 2− ) complicates the analysis and modelling of CO2 degassing, as the stripping process involves both mass transfer and chemical equilibria reactions. The two most pertinent considerations are (1) the relatively slow dehydroxylation of HCO3 − to CO2 , and (2) the ionization fractions (or relative proportions) of inorganic carbon species. The carbonate system is unable to instantaneously achieve equilibrium during the degassing process because the dehydroxylation of HCO3 − to CO2(aq) is a slow process in terms of chemical equilibria, taking over 1 min to reach full equilibrium (Grace and Piedrahita, 1994, p. 231). The slow dehydroxylation rate means that as a body of water passes through a stripping unit (which often only takes a few seconds), there is effectively no replacement of the degassed CO2 fraction from the pool of HCO3 − . The CO2 fraction is not replenished until the water re-establishes equilibrium some time after leaving the degassing unit, and the amount of CO2 that is re-formed from the carbonate pool depends on the ionization fraction. In a system where the alkalinity remains a fixed property of the water (such as inside a CO2 stripper), the ionization fraction is a function of the concentration of total inorganic carbon (Ct ), temperature and salinity. As was demonstrated in the companion paper, the slow dehydroxylation rate coupled with salinity-specific ionization fractions can cause significant differences in CO2 stripping efficiency between fresh and saline water (Moran, 2010).
2.1. Overview This study investigated the CO2 degassing efficiency of an airlift at two salinities: 0 and 35‰ NaCl. The trials were carried out in April–May 2008 at a recirculating aquaculture system (RAS) research facility described by Moran (2010). The water temperature for all trials was 14.8 ◦ C ± 0.4 (maximum deviation). A first set of trials was run at 0‰ using water from the local municipal water source. A second set of trials was run at 35‰ by the addition of NaCl (99.6% purity) to the system. The alkalinity of the freshwater was 5.41 meq L−1 , and the saltwater 5.11 meq L−1 . In this study the concentration of the unionized aqueous species (H2 CO3 (aq) and CO2 (aq)) have been summed into a single hypothetical species (CO∗2 ) because of the low hydration of CO2 (aq) (Stumm and Morgan, 1996, p. 97). 2.2. Configuration and testing of air-lift An air-lift was constructed with similar specifications to that used by Loyless and Malone (1998). A key difference was that in the present study the water from the air-lift discharged into a separate tank from which it was drawn (Fig. 1), whereas in the Loyless and Malone (1998) study the water discharged into the same tank from which it was drawn. By discharging the water into a second tank this study was able to calculate an absolute degassing efficiency, whereas the test apparatus of Loyless and Malone precluded the possibility of calculating this because the influent and effluent waters were mixed. The vertical riser of the air-lift was constructed of two pieces of 50 mm (OD) PVC pipe joined by a union located near the water surface. The head piece consisted of a 90◦ elbow attached to a 32 cm piece of straight pipe, followed by a 45◦ bend with a short piece of straight pipe which housed a CO2 probe (Fig. 1). The lift height was defined as the distance from the water surface to the centre line of the discharge pipe, and could be adjusted by inserting pieces of pipe of varying lengths into the union of the vertical riser section. Three lift heights were tested: 11, 16 and 25 cm. The air-lift was submerged 95 cm below the water surface of the test tank, and air was injected at 90 cm depth, as shown in Fig. 1. The airlift was placed in a test tank (112 cm (h) × 100 cm × 100 cm) with an adjustable water supply and standpipe system, and a water volume of 1020 L. In order to ensure the same head pressure in the test tank between different air-lift discharge rates, it was necessary to vary the incoming water flow rate so that there was approximately
Fig. 1. Experimental apparatus used to test air-lift stripping efficiency. The air line supplying the air-lift was monitored using pressure gauges (PG), a flowmeter (FM) and thermometer. Carbon dioxide concentration in the test tank and air-lift discharge was measured using OxyGuard CO2 Analyzers. Water discharged by the air-lift flowed into a second tank, and a floating water level indicator was used to measure the tank filling rate to derive Qw . The water level in the air-lift test tank was kept constant by adjusting the incoming water flow.
122
D. Moran / Aquacultural Engineering 43 (2010) 120–127
the same water flow rate out of the standpipe regardless of air-lift discharge rate. Water from the air-lift was discharged into another tank of the same dimensions, and both tanks drained into a sump of the RAS. Gaseous CO2 was added to the test tank via a 0–1 L min−1 flowmeter and fine pore diffuser placed in the bottom of the tank. The radial mixing caused by the inflowing water was sufficient to give a homogenous distribution of water and CO2 throughout the test tank. Airflow to the air-lift was supplied via a 4.8 kW Rietschle SAH 235 blower (Gardner Denver Schopfheim GmbH, Schopfheim, Germany), which was specified to deliver 4000 L min−1 at 38 kPa pressure difference. A 25 mm (OD) flexible hose ran from the air blower (located outside the RAS building) to the pressure (0–1 bar) and airflow indicator section of the air-lift set up (Fig. 1). Airflow to the air-lift was controlled via a ball valve, and was metered using a 20–200 nominal L min−1 Yokogawa rotameter (model RAQN, Roto Yokogawa Gmbh & Co., Wehr, Germany), which had a specified accuracy of 11–3% of the reading from the low to high range of the rotameter. Pressure gauges (0–1 bar) were installed both upstream and downstream of the flowmeter, and a thermometer was installed upstream of the flowmeter. A 25 mm hose ran from the meter section to the air-lift and connected to a tapping saddle on the airlift via a threaded hose barb. The air entered the air-lift through the external hose barb, so the injection method can be described as open-tube injection. Water flow rate was measured by recording the filling time of the discharge tank. Water level was determined using a floating sight level indicator located inside a pipe inside the discharge tank, which buffered the level indicator from the disturbance of the surface water (Fig. 1). Paired airflow and water discharge data were recorded for the saltwater trial to establish the relationship between Qg (gas flow rate) and Qw (airlift water discharge rate), but only Qw data was collected for the freshwater test run due to an accident which damaged the airflow rotameter. Carbon dioxide concentration of both the test tank and the airlift discharge water was measured using OxyGuard CO2 Analyzers (OxyGuard International A/S, Birkerød, Denmark). These gas analyzers have a submersible probe to measure the partial pressure of CO2 (pCO2 ) in water via infrared absorption. The pCO2 is converted to a temperature and salinity specific concentration via a solubility factor determined from the calibration procedure and a thermistor. The probes were calibrated according to the recommendations given by Moran et al. (2010). A key factor in determining the reaction time of the analyzer is the water velocity across the gas permeable probe membrane (Moran et al., (2010). The CO2 probe in the test tank was placed near the bottom of the tank, and was fitted with a 40 mm PVC plumbing union into which the outlet of a small submersible aquarium pump (5 W Eheim Compact 1000, 5 L min−1 , Eheim GmbH, Deizisau, Germany) was located, so that a constant stream of water passed over the probe membrane at a velocity of 40–47 cm s−1 . The probe of the CO2 analyzer measuring the air-lift discharge was set into the pipe in such a way that it had a constant high velocity stream of water (>45 cm s−1 ) passing over the membrane. Tests were undertaken to ensure that the probes measuring the influent and effluent CO2 concentration had adequate water velocities across the membranes to ensure comparable reaction times and measurements (Moran et al., (2010). The placement of the effluent CO2 probe very close to the outlet of the air-lift meant that it was recording the non-equilibrium CO2 concentration after degassing. The probes were calibrated according the recommendations given by Moran et al. (2010). To initiate a degassing trial the water flow rate was set to a desired level by a process of adjusting the airflow rate and measuring the water discharge rate. After the desired water discharge rate had been achieved, the water flow rate into the test tank was adjusted to ensure sufficient replenishment. The CO2 gas was then
Fig. 2. Air-lift pumping analysis using saline water. (a) Comparison of water discharge rate (Qw ) versus airflow rate (Qg ) at different lift heights. The lift heights (cm) are inset into the trend lines. The solid lines represent the data from the present study, while the dashed lines represent the data from Loyless and Malone (1998) for an air-lift of similar dimensions. (b) Comparison of Qw versus the gas:liquid ratio at different lift heights. The functions that describe the parabolic relationship are given in the below the figure. (c) Comparison of pumping efficiency (as defined by Nicklin, 1963) versus Qw .
dosed for as long as it took to reach between 50 and 80 mg L−1 in the test tank, at which point gassing was stopped. The CO2 concentration displayed on the OxyGuard CO2 Analyzers were recorded every 1–2 min until the air-lift discharge concentration decreased to less than 10 mg L−1 , which took 20–70 min depending on Qw . The Qw was measured five times during the degassing trial by repeatedly
D. Moran / Aquacultural Engineering 43 (2010) 120–127
123
Fig. 3. Comparison of air-lift CO2 mass transfer versus Qw at different salinities and lift heights. The lift heights are shown at the top centre of each graph. A trend line is fitted to the combined salinity data sets. The functions for log deficit ratio data are as follows: lift height 11 cm: y = 0.311 − 0.006x + 5.79E−5x2 , R2 = 0.702; lift height 16 cm: y = 0.439 − 0.009x + 8.083E−5x2 , R2 = 0.834; lift height 25 cm: y = 0.64 − 0.012x + 1.065E−4x2 , R2 = 0.926. The functions for (kL a)20 are as follows: lift height 11 cm: y = 0.016(0.028x) , R2 = 0.90; lift height 16 cm: y = 0.018(0.027x) , R2 = 0.96; lift height 25 cm: y = 0.035(0.023x) , R2 = 0.99.
emptying and partially refilling (150–400 L) the discharge tank, and recording the length of time it took to refill. The volume of water discharged was calculated on the basis of tank area and refill height. Refill height was measured using a water level indicator located inside a pipe within the discharge tank (Fig. 1). 2.3. Mass transfer and stripping efficiency calculations The performance of the air-lift was analyzed in two ways. The first was the mass transfer coefficient (kL a), which is described as eq
kL a =
eq
non-eq
ln(CO2(air) − CO2(in) /CO2(air) − CO2(eff ) ) t
,
(1) eq
where kL a = mass transfer coefficient, 1/time, CO2(air) = CO∗2 concentration of water in equilibrium with the air entering the air-lift, CO2(in) = CO∗2 concentration of influent water, non-eq ∗ CO2(eff ) = CO2 concentration of effluent water prior to kinetic equilibration reactions, t = time. The second measure of stripping efficiency was the CO2 stripping efficiency (% CO2 stripped in a single pass), which is defined according to Grace and Piedrahita (1993, p. 501) as CO2 stripping efficency (%) =
CO2 − CO2 × 100, CO2 − CO2
(2)
eq
CO2(eff ) = CO∗2 where CO2(in) = influent CO∗2 concentration, concentration of effluent water after equilibration reactions, eq CO2(eff ) = CO∗2 concentration of water in equilibrium with the injected air. Carbon dioxide stripping efficiency differs conceptually from kL a in that uses the CO∗2 concentration of the effluent water after equieq librium reactions (CO2(eff ) ) to calculate how effectively degassing proceeded. Carbon dioxide stripping efficiency therefore combines the mass transfer process and chemical kinetic equilibria. The methodology used to derive kL a and CO2 stripping efficiencies from the recorded CO2 concentration data is described in detail in Moran (2010), with some minor differences in kL a calculations due to the nature of the degassing methodology (i.e. packing height is used to standardise kL a data for cascade columns, whereas time is used to standardise air-lift data). A general description of the calcueq lation of kL a is as follows. First, a plot was made of CO2(air) − CO2(in) eq
non-eq
versus CO2(air) − CO2(eff ) (i.e. the numerator versus denominator of the log deficit ratio term of Eq. (1)), and the slope of the regression used to calculate the log deficit ratio (R2 values for regressions eq >0.98). The calculation of CO2(air) requires knowledge of the CO2 concentration of air being injected into the air-lift. In the current study a value of 380 ppm CO2 was used based on current atmospheric observations from the NOAA Mauna Loa Observatory, Hawaii. The solubility constants for CO2 in fresh and saline water
124
D. Moran / Aquacultural Engineering 43 (2010) 120–127
were calculated from the equations given by Weiss (1974). To complete the calculation of kL a the residence time of water within the air-lift was needed to account for the time component of Eq. (1) (Reinemann and Timmons, 1989). The volume of the air-lift was calculated for the different lift heights (2.21, 2.29 and 2.44 L at 11, 16, and 25 cm lift height, respectively), allowing the residence time to be derived by dividing air-lift volume (L) by Qw (L s−1 ). The residence time (s) was used to calculate kL a, with units of s−1 . Finally, kL a was converted to a standard reference temperature of 20 ◦ C ((kL a)20 ) using the Arrhenius temperature relation kL a = (kL a)20 Temp−20
(3)
where is the temperature correction factor (=1.024), and Temp is the water temperature (◦ C) (Sincero and Sincero, 2003). The provision of kL a data allows readers to calculate the predicted airlift CO2 stripping efficiency for waters of different influent CO2 concentrations, salinities, temperatures and alkalinities using the carbonate chemistry approach outlined in Moran (2010). Unlike kL a, CO2 stripping efficiency varies according to influent CO∗2 concentration, therefore, CO2 stripping efficiency was calculated at two reference CO2(in) concentrations, 10 and 40 mg L−1 . As CO2 stripping efficiency can only be derived from the CO∗2 conceneq tration of the effluent water after equilibrium reactions (CO2(eff ) , which was not measured), this study used the same approach of the companion paper (Moran, 2010) to inter-convert CO∗2 concentration and total inorganic carbon concentration at equilibrium. The principle was to: (a) calculate the influent total inorganic carbon (Ct(inf) ) concentration based on the CO∗2 concentration and alkalinity; (b) calculate the mass of CO2 removed during degassing using the kL a mass transfer data; (c) derive the effluent total inorganic eq carbon (Ct(eff) ) concentration; (d) calculate CO2(eff ) from Ct(eff) using knowledge of the carbonate equilibrium relationship; and (e) use Eq. (2) to calculate the CO2 stripping efficiency. To complete part (a) Eqs. (13) and (14) of Moran (2010) were used to derive Ct(inf) from CO2(in) . Note that the water temperature, alkalinities and salinities were the same in both studies, thereby allowing for the use of the Ct -CO∗2 conversion functions given in Moran (2010). In part (b) the non-eq expected value of CO2(eff ) was calculated for a given value of CO2(in) eq
using Eq. (1), as the variables kL a, CO2(air) , CO2(in) , and t (residence time) were known for each salinity, lift height and Qw combination. non-eq The mass of CO2 degassed could be obtained via CO2(in) − CO2(eff ) . The results of parts (a) and (b) could then be used to calculate part (c), namely Ct(eff) . In part (d), Eqs. (16) and (17) of Moran (2010) were used to obtain an equivalent CO∗2 concentration at equilibrium eq for Ct(eff) . Finally, in part (e), the CO2(eff ) values from part (d) could eq
be used along with CO2(in) and CO2(air) to calculate CO2 stripping efficiency using Eq. (2). 2.4. Analysis of air-lift data The airflow recorded from the rotameter was corrected to standard temperature pressure using the ideal gas law, with an assumed compressibility factor of Z = 1 given the relatively low pressures used. The data for Qw and Qg were then plotted on a graph along with the logarithmic curves described by Loyless and Malone (1998) for other lift heights. The relationship between Qw and Qg in the present study was not suitably described by the logarithmic function of Loyless and Malone (1998), so a number of other functions were tested using the curve fitting software TableCurve (v4, Jandel Scientific Software, CA, U.S.A.). The only functions with a high correlation coefficient were too over-parameterised to be useful, so no function was fitted. The gas:liquid ratio (G:L) was plotted against Qw to help interpret the importance of relative gas volume in mass transfer. The pumping efficiency was evaluated using the
Fig. 4. Relationship between log deficit ratio (LDR) and three key airlift variables: lift height, pumping efficiency (G:L ratio) and water discharge rate (Qw ). (a) Dependence of LDR on lift height and G:L ratio. LDR = 0.018 + 0.010 lift + 0.111 G:L, R2 = 0.87. (b) Dependence of LDR on lift height and Qw LDR = 0.462 − 0.014 lift − 0.009 Qw + 7.23 × 10−4 lift2 + 7.82 × 10−5 Qw 2 , R2 = 0.85.
Nicklin (1963) efficiency equation cited in Reinemann et al. (1990, Eq. 14), which is valid to within 1% for air-lifts with less than 5 m submergence and a diameter greater than 20 mm. Air-lift pumping efficiency (, dimensionless) is defined as the net work done lifting the liquid divided by the work the air does as it expands isothermally Nicklin (1963). Both the log deficit ratio and the (kL a)20 were used as measures of the mass transfer performance of the air-lift. The log deficit ratio represents the proportion of CO2 transferred in the time it takes for water to travel from the inlet to the outlet of the air-lift. The (kL a)20 is the log deficit ratio divided by the residence time of water within the air-lift. The former measure of mass transfer is useful for understanding the general efficiency of an air-lift, while the latter is a standard unit used in the engineering literature and cross-study comparisons. The mass transfer performance results were plotted against Qw for each lift height, and an appropriate trend curve fitted to the combined salinity data set. The CO2 stripping efficiency data were plotted at two reference influent CO∗2 concentrations, 10 and 40 mg L−1 . The CO2 degassing data from the current study was converted to a format that could be compared to the Loyless and Malone (1998) data, which to the author’s knowledge is the only other published study on the CO2 stripping efficiency of an air-lift. Loyless and Malone (1998) published CO2 degassing data for a lift height of 15.24 cm. The (kL a)20 values were not given by Loyless and Malone (1998), but rather the degassing efficiency was expressed as the standard CO2 transfer rate (SCTR, g CO2 h−1 ). SCTR = (kL a)20 [Cm − (Cs )20 ]V 60 × 10−3
(4)
D. Moran / Aquacultural Engineering 43 (2010) 120–127
125
Fig. 5. Comparison of CO2 stripping efficiency versus Qw at different salinities and lift heights. The stripping efficiency at 15 ◦ C is given for two influent CO2 concentrations, 40 and 10 mg L−1 .The lift heights are shown at the top centre of each graph.
where (Cs )20 is saturation concentration at 20 ◦ C (0.5 mg L−1 ), Cm is standard measured concentration (arbitrarily defined as 1 mg L−1 ), V is volume of water (L), 60 is conversion from min to h, and 10−3 converts from mg to g. The SCTR was calculated for the air-lift data of the current study using the same values of (Cs )20 and Cm in order to allow for direct comparison. The SCTR data of Loyless and Malone (1998) was expressed as a function of Qw and plotted alongside the results from the current study.
3. Results and discussion The rate of water discharge from the air-lift increased in a nonlinear manner with increasing airflow (Fig. 2a), though the function that best explained the relationship was indeterminable. The relationship between Qg and Qw observed in the present study was different to that reported by Loyless and Malone (1998), whose data showed a clear logarithmic correlation for an air-lift of almost identical proportions. In the current study, there was a rapid increase in Qw from the lowest gas flow rate of 20 L min−1 , followed by a more curvilinear increase in Qw after Qg reached 30 L min−1 (Fig. 2a). It appeared that 20–30 L min−1 represented a threshold Qg that was required to initiate water flow past the level of the horizontal discharge section. This was possibly an artefact of the flowmeter used in the present study, which was only specified for a range of 20–200 L min−1 . Loyless and Malone (1998) used three airflow meters with differing operating ranges, so were probably able to obtain more accurate estimates of Qg at low airflow rates (i.e. <20 L min−1 ).
Increasing lift height reduced Qw , but there was disagreement with the relative positioning of the lift height curves between the present study and that of Loyless and Malone (1998) (Fig. 2a). In the present study, the values of Qw at each lift height tended to be higher for a given value of Qg above 30 L min−1 than that given by Loyless and Malone (1998). The salinity of the water used for the pumping efficiency measurements in the current study (35‰ NaCl) differed from the freshwater used by Loyless and Malone (1998), and the lack of comparative studies make it difficult to precisely determine the influence of salinity on air-lift pumping performance. There was consensus between the two studies that at lift heights above 11 cm it appears impossible to achieve water flow rates above c.110 L min−1 for the specified air-lift. There was a positive parabolic relationship between G:L ratio versus Qw (Fig. 2b), and a negative parabolic relationship between pumping efficiency and Qw (Fig. 2c). The G:L ratio was high (∼1.9–2.0) at low Qw and represented the initial input energy required to raise the water up the vertical riser section to the discharge pipe. The air-lift increased in pumping efficiency rapidly thereafter, and was at maximum efficiency at 50–70 L min−1 water flow (Fig. 2c). After this point the G:L ratio increased with Qw and the pumping efficiency decreased. The pumping efficiency at the minimum and maximum water flow rates were 25–40% lower than the maximum recorded efficiencies (Fig. 2c). Salinity had no obvious or consistent effect on the CO2 mass transfer characteristics of the air-lift (i.e. log deficit ratio and (kL a)20 ), with data points distributed evenly about the combined data curves (Fig. 3). There was some evidence of a higher mass transfer efficiency in freshwater at high aeration rates (i.e.
126
D. Moran / Aquacultural Engineering 43 (2010) 120–127
>100 L min−1 Qw , Fig. 3), however, there were not enough data points in this flow range to objectively test this trend. Ruen-ngam et al. (2008) investigated the effect of salinity on bubble size and oxygen mass transfer in an air-lift contactor (bubble column), and found that increasing seawater salinity resulted in smaller bubble diameters and a higher specific surface area for mass transfer. This finding intuitively suggests that mass transfer should proceed more efficiently in saltwater compared to freshwater. However, Ruen-ngam et al. (2008) only found differences in (kL a)20 between salinities at the highest aeration rates, and at the highest aeration rates mass transfer was greater in freshwater. Further testing of the air-lift in the current study at high aeration rates may yield similar salinity effects on mass transfer as that reported by Ruen-ngam et al. (2008). The concentration driving force varied with Qw in a parabolic manner, such that it initially decreased with Qw to a minimum at approximately 50 L min−1 , before increasing again (Fig. 3). The concentration driving force at the low and high end of the Qw range was approximately 75–100% greater than that at the medium Qw range (40–60 L min−1 ). This general relationship was the same for all lift heights and salinities. The parabolic relationship between log deficit ratio and Qw can be traced to the G:L ratio the air-lift was operating at. When the G:L ratio was increased, there was more air present per unit water to equalize the CO2 concentration gradient, thereby leading to greater mass transfer. The dependence of CO2 mass transfer on the G:L ratio was evident in Fig. 4a, as was the effect of lift height on CO2 mass transfer. Increasing lift height resulted in increased CO2 mass transfer at all water flow rates. The likely explanation for this is that the total contact time between the air and water increases with lift height, thereby allowing for more complete CO2 mass transfer across the concentration gradient. Comparison of the coefficients that describe log deficit ratio as a function of lift height and G:L ratio (Fig. 4a) indicated that a 5 cm increase in lift height was approximately equal to G:L increase of 0.5. Fig. 4b shows the relationship between Qw , lift height and log deficit ratio should a reader wish to calculate the CO2 mass transfer coefficient for a given Qw and lift height. The (kL a)20 values increased exponentially with Qw (Fig. 3), implying that mass transfer proceeded more rapidly as the volume of injected gas increased. In accordance with the log deficit ratio results, the CO2 stripping efficiency changed with Qw in a parabolic manner, and increasing lift height increased CO2 stripping efficiency (Fig. 5). Salinity had little effect on CO2 stripping efficiency, however, CO2 stripping efficiency was slightly higher at elevated influent CO2 concentrations. At an influent concentration of 40 mg L−1 the CO2 stripping efficiency was 1–3% higher than at an influent of 10 mg L−1 (Fig. 5). The relatively minimal effect of water salinity and influent CO∗2 concentration on the CO2 stripping efficiency of the air-lift stands in contrast to the significant effects these variables had on the stripping efficiency of a cascade column in a companion study (Moran, 2010). The reason for the discrepancy in the effects of salinity and influent CO∗2 concentration lay in the magnitude of the mass transfer efficiencies recorded in each study. The log deficit ratio values recorded for the cascade column tested by (Moran, 2010) were in the range of 1.8–2.3, which were much higher than the values measured for the air-lift in the current study (0.18–0.50, Fig. 3). The higher log deficit ratio of the cascade column represents increased mass transfer of CO2 , which considerably decreases the Ct and H+ concentration of the effluent water. The sizeable decrease in Ct and H+ concentration means that the effluent water has markedly different ionization fractions of inorganic carbon species compared to the influent water. Salinity-dependent differences in the ionization fractions of inorganic carbon species are greatest when the Ct concentration is low (relative to the alkalinity, Fig. 1 of Moran, 2010). The result is that for any given influent CO∗2 concentration,
Fig. 6. Comparison of standard CO2 transfer rates (SCTR) given by Loyless and Malone (1998) and the current study. The air-lift of Loyless and Malone (1998) was set to a lift height of 15.2 cm and was nearly identical in construction to that used in the current study. The function used to represent the data of Loyless and Malone (1998) is SCTR = 0.25 + 0.007 exp((Qw + 89)/40).
waters of different salinities exit a highly efficient degasser with different post-equilibrium effluent CO∗2 concentrations. However, in a less efficient degasser there is only a moderate decrease in Ct and H+ concentration, therefore, the ionization fractions of the influent and effluent water only differ a little, and influent CO∗2 concentration and salinity have less of an effect on CO2 stripping efficiency. The standard CO2 transfer rates increased exponentially with Qw (Fig. 6). The SCTR trend curves for lift heights 11 and 16 cm were similar, while at 25 cm lift height there was an increased SCTR for any given value of Qw (Fig. 6). The SCTR data from Loyless and Malone (1998) for a nearly identical air-lift at 15 cm lift height did not correlate closely with the data of the current study at lowmedium values of Qw . The Loyless and Malone (1998) trend line gave substantially higher SCTR values for water flow rates below 90 L min−1 (Fig. 6). The reason for the discrepancy between the studies is unclear, as Loyless and Malone (1998) did not report (kL a)20 values (upon which SCTR is reliant), nor the method used to calculate the mass of CO2 degassed.
D. Moran / Aquacultural Engineering 43 (2010) 120–127
127
4. Conclusions
References
There were significant differences in the pumping and CO2 degassing characteristics of the air-lift in the present paper compared to that of Loyless and Malone (1998), despite both studies using air-lifts of almost identical construction. The reasons for the differences are unclear and warrant further investigation for the purpose of consensus building. Salinity was not found to have a significant effect on the CO2 mass transfer characteristics of the airlift tested. In contrast to a companion study of a cascade column, the CO2 stripping efficiency of the air-lift did not differ with salinity, which was attributed to the comparatively poor mass transfer performance of the air-lift. The pumping efficiency of the air-lift was highest when Qw was between 60 and 70 L min−1 , whereas the absolute rate of CO2 transfer increased exponentially with Qw . A general equation was derived for the airlift used in this study that allows one to calculate the mass transfer coefficient for a given lift height, Qw , or G:L ratio. The mass transfer coefficient can then be used to calculate the CO2 stripping efficiency for any water type (i.e. temperature, alkalinity, salinity and influent CO2 concentration).
Grace, G.R., Piedrahita, R.H., 1993. Carbon dioxide control with a packed column aerator. In: Wang, J.K. (Ed.), Techniques for Modern Aquaculture. American Society of Agricultural Engineers, Saint Joseph, MI, pp. 496–505. Grace, G.R., Piedrahita, R.H., 1994. Carbon dioxide control. In: Timmons, M.B., Losordo, T.M. (Eds.), Aquaculture Water Reuse Systems: Engineering Design and Management. Elsevier, New York, USA, pp. 209–234. La Motta, E.J., 1995. Chemical analysis of CO2 removal in tray aerators. J. Am. Water Resour. Assoc. 31, 207–216. Loyless, J.C., Malone, R.F., 1998. Evaluation of air-lift pump capabilities for water delivery, aeration, and degasification for application to recirculating aquaculture systems. Aquacult. Eng. 18, 117–133. Moran, D., 2010. Carbon dioxide degassing in fresh and saline water. I. Degassing performance of a cascade column. Aquacult. Eng. 43, 29–36. Moran, D., Tirsgård, B., Steffensen, J.F., 2010. The accuracy and limitations of a new meter used to measure aqueous carbon dioxide concentration. Aquacult. Eng. 43, 101–107. Nicklin, D.J., 1963. The air-lift pump: theory and optimization. Trans. Inst. Chem. Eng.. Reinemann, D.J., Timmons, M.B., 1989. Prediction of oxygen transfer and total dissolved gas pressure in airlift pumping. Aquacult. Eng. 8, 29–46. Reinemann, D.J., Parlange, J.Y., Timmons, M.B., 1990. Theory of small-diameter airlift pumps. Int. J. Multiphase Flow 16, 113–122. Ruen-ngam, D., Wongsuchoto, P., Limpanuphap, A., Charinpanitkul, T., Pavasant, P., 2008. Influence of salinity on bubble size distribution and gas–liquid mass transfer in airlift contactors. Chem. Eng. J. 141, 222–232. Sincero, A.P., Sincero, G.A., 2003. Chapter 9: Aeration, Absorption and Stripping, Physical–Chemical Treatment of Water and Wastewater. CRC Press, Florida, USA. Stumm, W., Morgan, J.J., 1996. Aquatic Chemistry: Chemical Equilibria and Rates in Natural Waters. Wiley-Interscience, New York, USA. Summerfelt, S.T., Vinci, B.J., Piedrahita, R.H., 2000. Oxygenation and carbon dioxide control in water reuse systems. Aquacult. Eng. 22, 87–108. Summerfelt, S.T., Davidson, J., Waldrop, T., 2003. Evaluation of full-scale carbon dioxide stripping columns in a coldwater recirculating system. Aquacult. Eng. 28, 155–169. Timmons, M.B., Ebeling, J.M., Wheaton, F.W., Summerfelt, S.T., Vinci, B.J., 2002. Recirculating Aquaculture Systems. Cayuga Aquaculture Ventures, Ithaca, NY, 769 pp. Watten, B.J., Sibrell, P.L., Montgomery, G.A., Tsukuda, S.M., 2004. Modification of pure oxygen absorption equipment for concurrent stripping of carbon dioxide. Aquacult. Eng. 32, 183–208. Weiss, R.F., 1974. Carbon dioxide in water and seawater: the solubility of a non-ideal gas. Mar. Chem. 2, 203–215.
Acknowledgements I am thankful for the technical and practical support given by M. Fülberth, and logistical support given by J. Bregnballe, H. Jarlbæk, B.H Olsen, and J.G. Støttrup. I also thank the three anonymous reviewers for reviewing and improving this manuscript. This work was partially funded by the EU Fisheries Sector Program FIUF and the Danish Ministry of Food, Agriculture and Fisheries. The author was supported by a fellowship from the New Zealand Foundation for Research Science and Technology.
Contents lists available at ScienceDirect
Aquacultural Engineering journal homepage: www.elsevier.com/locate/aqua-online
Carbon dioxide degassing in fresh and saline water. II: Degassing performance of an air-lift Damian Moran ∗ DTU Aqua, Technical University of Denmark, Charlottenlund, Denmark
a r t i c l e
i n f o
Article history: Received 8 June 2010 Accepted 16 September 2010 Keywords: CO2 stripping Degassing Recirculating Reuse Salinity
a b s t r a c t A study was undertaken to measure the efficiency with which carbon dioxide was stripped from freshwater (0‰) and saline water (35‰ NaCl) passing through an air-lift at 15 ◦ C. The air-lift was constructed of 50 mm (OD) PVC pipe submerged 95 cm in a tank, had an adjustable air injection rate, and could be adjusted to three lifting heights: 11, 16 and 25 cm. The gas to liquid ratio (G:L) was high (∼1.9–2.0) at low water discharge rates (Qw ) and represented the initial input energy required to raise the water up the vertical riser section to the discharge pipe. The air-lift increased in pumping efficiency rapidly thereafter, to a G:L minima of 0.3–0.6 at 60–70 L min−1 . After this point the G:L ratio increased with Qw , representing decreasing air-lift pumping efficiency. The CO2 concentration of the influent and effluent water was measured using submersible infrared CO2 probes over a range of influent CO2 concentrations. The CO2 mass transfer coefficient [(kL a)20 ] ranged from 0.025 to 0.468. Increasing lift height increased mass transfer, which was attributed to both the increased G:L ratio and the contact time inside the air-lift. The relative effect of lift height and pumping rate on mass transfer was such that a 5 cm increase in lift height was approximately equal to a G:L increase of 0.5. The CO2 stripping efficiency was effectively the same between salinities, and the influent CO2 concentration only had a modest effect on CO2 stripping efficiency. At an influent concentration of 40 mg L−1 the CO2 stripping efficiency was 1–3% higher than at an influent of 10 mg L−1 . The relatively minor effects of salinity and influent CO2 concentration on stripping efficiency contrasted with a companion study investigating the stripping efficiency of a cascade column. The difference was attributed to the low-to-moderate mass transfer efficiencies of the air-lift. A general equation was derived for the airlift that allows one to calculate the mass transfer coefficient for a given lift height, Qw , or G:L ratio. The mass transfer coefficient can then be used to calculate the CO2 stripping efficiency for any water type (i.e. temperature, alkalinity, salinity and influent CO2 concentration). © 2010 Elsevier B.V. All rights reserved.
1. Introduction To design a system that meets the CO2 water quality requirements for a species, there needs to be knowledge about the biological tolerance of stock to CO2 , and the influx and efflux of inorganic carbon. This allows for a mass-balance analysis of the dynamics of inorganic carbon in the system to be carried out with a target operating CO2 concentration in mind. The aim of the current study was to provide data on the efflux of CO2 from an air-lift, a device used in aquaculture to pump water, ingas O2 and degas CO2 . Reinemann and Timmons (1989) and Timmons et al. (2002) provide an introduction to the theory and operation of air-lifts, but essentially an air-lift is a pipe submerged vertically into a water column with a 90◦ bend below, at or above the water surface. When
∗ Present address: Department of Biology, Lund University, Sölvegatan 35, S-22362 Lund, Sweden. Tel.: +46 46 22 27 785; fax: +46 46 22 24 425. E-mail address: [email protected]. 0144-8609/$ – see front matter © 2010 Elsevier B.V. All rights reserved. doi:10.1016/j.aquaeng.2010.09.001
air is injected at the bottom of pipe it rises and draws water with it, eventually discharging after rounding the 90◦ bend. To the author’s knowledge there is only one published study evaluating the CO2 degassing performance of an air-lift. Loyless and Malone (1998) measured the rate at which CO2 was removed from a tank of freshwater with an air-lift operating inside. While the CO2 removal rate was specific to the air-lift/tank configuration tested and could not be converted to a generalised measure of degassing efficiency (e.g. mass transfer or CO2 stripping efficiency), the study did show how lift height and air injection rate affected relative degassing performance. The current study adds to the work of Loyless and Malone (1998) using an almost identical air-lift to directly measure CO2 degassing performance, thereby allowing researchers to easily combine the lifting and pumping efficiency data of Loyless and Malone (1998) with the CO2 degassing efficiency data from this paper. A novel aspect of the current paper is the comparison of CO2 degassing efficiency in both fresh (0‰) and saline water (35‰ NaCl). Most other studies published to date on CO2 degassing have been solely concerned with freshwater (Grace
D. Moran / Aquacultural Engineering 43 (2010) 120–127
121
2. Materials and methods Nomenclature Alkt CO∗2 eq CO2(air)
CO2(in) non-eq CO2(eff ) eq
CO2(eff ) Ct Ct(in) Ct(eff) G:L kL a pCO2 Qg Qw t
total alkalinity sum of the unionized aqueous species H2 CO3 (aq) and CO2 (aq) CO∗2 concentration of water in equilibrium with air entering the air-lift CO∗2 concentration of influent water CO∗2 concentration of effluent water prior to equilibration reactions CO∗2 concentration of effluent water after equilibration reactions total inorganic carbon Ct concentration of influent water Ct concentration of effluent water gas:liquid ratio mass transfer coefficient (s−1 residence time) partial pressure of CO2 gas injection rate (L min−1 ) water flow rate (L min−1 ) time (s)
and Piedrahita, 1993, 1994; La Motta, 1995; Summerfelt et al., 2000, 2003; Watten et al., 2004).
1.1. Carbonate chemistry during CO2 degassing Grace and Piedrahita (1994) provide an extensive introduction to the gas transfer and chemistry processes that occur during CO2 degassing, and a companion paper describes how salinity influences CO2 degassing (Moran, 2010). A summary of the CO2 degassing process is as follows. As a body of water passes through a stripping unit, CO2 diffuses from the water into the air along the concentration gradient, a process known as mass transfer. The important difference between CO2 and other gases is that when CO2 dissolves in water, it undergoes chemical reactions, forming carbonates and H+ . The dissolution of CO2 into a number of aqueous species (CO2 (aq), H2 CO3 , HCO3 − and CO3 2− ) complicates the analysis and modelling of CO2 degassing, as the stripping process involves both mass transfer and chemical equilibria reactions. The two most pertinent considerations are (1) the relatively slow dehydroxylation of HCO3 − to CO2 , and (2) the ionization fractions (or relative proportions) of inorganic carbon species. The carbonate system is unable to instantaneously achieve equilibrium during the degassing process because the dehydroxylation of HCO3 − to CO2(aq) is a slow process in terms of chemical equilibria, taking over 1 min to reach full equilibrium (Grace and Piedrahita, 1994, p. 231). The slow dehydroxylation rate means that as a body of water passes through a stripping unit (which often only takes a few seconds), there is effectively no replacement of the degassed CO2 fraction from the pool of HCO3 − . The CO2 fraction is not replenished until the water re-establishes equilibrium some time after leaving the degassing unit, and the amount of CO2 that is re-formed from the carbonate pool depends on the ionization fraction. In a system where the alkalinity remains a fixed property of the water (such as inside a CO2 stripper), the ionization fraction is a function of the concentration of total inorganic carbon (Ct ), temperature and salinity. As was demonstrated in the companion paper, the slow dehydroxylation rate coupled with salinity-specific ionization fractions can cause significant differences in CO2 stripping efficiency between fresh and saline water (Moran, 2010).
2.1. Overview This study investigated the CO2 degassing efficiency of an airlift at two salinities: 0 and 35‰ NaCl. The trials were carried out in April–May 2008 at a recirculating aquaculture system (RAS) research facility described by Moran (2010). The water temperature for all trials was 14.8 ◦ C ± 0.4 (maximum deviation). A first set of trials was run at 0‰ using water from the local municipal water source. A second set of trials was run at 35‰ by the addition of NaCl (99.6% purity) to the system. The alkalinity of the freshwater was 5.41 meq L−1 , and the saltwater 5.11 meq L−1 . In this study the concentration of the unionized aqueous species (H2 CO3 (aq) and CO2 (aq)) have been summed into a single hypothetical species (CO∗2 ) because of the low hydration of CO2 (aq) (Stumm and Morgan, 1996, p. 97). 2.2. Configuration and testing of air-lift An air-lift was constructed with similar specifications to that used by Loyless and Malone (1998). A key difference was that in the present study the water from the air-lift discharged into a separate tank from which it was drawn (Fig. 1), whereas in the Loyless and Malone (1998) study the water discharged into the same tank from which it was drawn. By discharging the water into a second tank this study was able to calculate an absolute degassing efficiency, whereas the test apparatus of Loyless and Malone precluded the possibility of calculating this because the influent and effluent waters were mixed. The vertical riser of the air-lift was constructed of two pieces of 50 mm (OD) PVC pipe joined by a union located near the water surface. The head piece consisted of a 90◦ elbow attached to a 32 cm piece of straight pipe, followed by a 45◦ bend with a short piece of straight pipe which housed a CO2 probe (Fig. 1). The lift height was defined as the distance from the water surface to the centre line of the discharge pipe, and could be adjusted by inserting pieces of pipe of varying lengths into the union of the vertical riser section. Three lift heights were tested: 11, 16 and 25 cm. The air-lift was submerged 95 cm below the water surface of the test tank, and air was injected at 90 cm depth, as shown in Fig. 1. The airlift was placed in a test tank (112 cm (h) × 100 cm × 100 cm) with an adjustable water supply and standpipe system, and a water volume of 1020 L. In order to ensure the same head pressure in the test tank between different air-lift discharge rates, it was necessary to vary the incoming water flow rate so that there was approximately
Fig. 1. Experimental apparatus used to test air-lift stripping efficiency. The air line supplying the air-lift was monitored using pressure gauges (PG), a flowmeter (FM) and thermometer. Carbon dioxide concentration in the test tank and air-lift discharge was measured using OxyGuard CO2 Analyzers. Water discharged by the air-lift flowed into a second tank, and a floating water level indicator was used to measure the tank filling rate to derive Qw . The water level in the air-lift test tank was kept constant by adjusting the incoming water flow.
122
D. Moran / Aquacultural Engineering 43 (2010) 120–127
the same water flow rate out of the standpipe regardless of air-lift discharge rate. Water from the air-lift was discharged into another tank of the same dimensions, and both tanks drained into a sump of the RAS. Gaseous CO2 was added to the test tank via a 0–1 L min−1 flowmeter and fine pore diffuser placed in the bottom of the tank. The radial mixing caused by the inflowing water was sufficient to give a homogenous distribution of water and CO2 throughout the test tank. Airflow to the air-lift was supplied via a 4.8 kW Rietschle SAH 235 blower (Gardner Denver Schopfheim GmbH, Schopfheim, Germany), which was specified to deliver 4000 L min−1 at 38 kPa pressure difference. A 25 mm (OD) flexible hose ran from the air blower (located outside the RAS building) to the pressure (0–1 bar) and airflow indicator section of the air-lift set up (Fig. 1). Airflow to the air-lift was controlled via a ball valve, and was metered using a 20–200 nominal L min−1 Yokogawa rotameter (model RAQN, Roto Yokogawa Gmbh & Co., Wehr, Germany), which had a specified accuracy of 11–3% of the reading from the low to high range of the rotameter. Pressure gauges (0–1 bar) were installed both upstream and downstream of the flowmeter, and a thermometer was installed upstream of the flowmeter. A 25 mm hose ran from the meter section to the air-lift and connected to a tapping saddle on the airlift via a threaded hose barb. The air entered the air-lift through the external hose barb, so the injection method can be described as open-tube injection. Water flow rate was measured by recording the filling time of the discharge tank. Water level was determined using a floating sight level indicator located inside a pipe inside the discharge tank, which buffered the level indicator from the disturbance of the surface water (Fig. 1). Paired airflow and water discharge data were recorded for the saltwater trial to establish the relationship between Qg (gas flow rate) and Qw (airlift water discharge rate), but only Qw data was collected for the freshwater test run due to an accident which damaged the airflow rotameter. Carbon dioxide concentration of both the test tank and the airlift discharge water was measured using OxyGuard CO2 Analyzers (OxyGuard International A/S, Birkerød, Denmark). These gas analyzers have a submersible probe to measure the partial pressure of CO2 (pCO2 ) in water via infrared absorption. The pCO2 is converted to a temperature and salinity specific concentration via a solubility factor determined from the calibration procedure and a thermistor. The probes were calibrated according to the recommendations given by Moran et al. (2010). A key factor in determining the reaction time of the analyzer is the water velocity across the gas permeable probe membrane (Moran et al., (2010). The CO2 probe in the test tank was placed near the bottom of the tank, and was fitted with a 40 mm PVC plumbing union into which the outlet of a small submersible aquarium pump (5 W Eheim Compact 1000, 5 L min−1 , Eheim GmbH, Deizisau, Germany) was located, so that a constant stream of water passed over the probe membrane at a velocity of 40–47 cm s−1 . The probe of the CO2 analyzer measuring the air-lift discharge was set into the pipe in such a way that it had a constant high velocity stream of water (>45 cm s−1 ) passing over the membrane. Tests were undertaken to ensure that the probes measuring the influent and effluent CO2 concentration had adequate water velocities across the membranes to ensure comparable reaction times and measurements (Moran et al., (2010). The placement of the effluent CO2 probe very close to the outlet of the air-lift meant that it was recording the non-equilibrium CO2 concentration after degassing. The probes were calibrated according the recommendations given by Moran et al. (2010). To initiate a degassing trial the water flow rate was set to a desired level by a process of adjusting the airflow rate and measuring the water discharge rate. After the desired water discharge rate had been achieved, the water flow rate into the test tank was adjusted to ensure sufficient replenishment. The CO2 gas was then
Fig. 2. Air-lift pumping analysis using saline water. (a) Comparison of water discharge rate (Qw ) versus airflow rate (Qg ) at different lift heights. The lift heights (cm) are inset into the trend lines. The solid lines represent the data from the present study, while the dashed lines represent the data from Loyless and Malone (1998) for an air-lift of similar dimensions. (b) Comparison of Qw versus the gas:liquid ratio at different lift heights. The functions that describe the parabolic relationship are given in the below the figure. (c) Comparison of pumping efficiency (as defined by Nicklin, 1963) versus Qw .
dosed for as long as it took to reach between 50 and 80 mg L−1 in the test tank, at which point gassing was stopped. The CO2 concentration displayed on the OxyGuard CO2 Analyzers were recorded every 1–2 min until the air-lift discharge concentration decreased to less than 10 mg L−1 , which took 20–70 min depending on Qw . The Qw was measured five times during the degassing trial by repeatedly
D. Moran / Aquacultural Engineering 43 (2010) 120–127
123
Fig. 3. Comparison of air-lift CO2 mass transfer versus Qw at different salinities and lift heights. The lift heights are shown at the top centre of each graph. A trend line is fitted to the combined salinity data sets. The functions for log deficit ratio data are as follows: lift height 11 cm: y = 0.311 − 0.006x + 5.79E−5x2 , R2 = 0.702; lift height 16 cm: y = 0.439 − 0.009x + 8.083E−5x2 , R2 = 0.834; lift height 25 cm: y = 0.64 − 0.012x + 1.065E−4x2 , R2 = 0.926. The functions for (kL a)20 are as follows: lift height 11 cm: y = 0.016(0.028x) , R2 = 0.90; lift height 16 cm: y = 0.018(0.027x) , R2 = 0.96; lift height 25 cm: y = 0.035(0.023x) , R2 = 0.99.
emptying and partially refilling (150–400 L) the discharge tank, and recording the length of time it took to refill. The volume of water discharged was calculated on the basis of tank area and refill height. Refill height was measured using a water level indicator located inside a pipe within the discharge tank (Fig. 1). 2.3. Mass transfer and stripping efficiency calculations The performance of the air-lift was analyzed in two ways. The first was the mass transfer coefficient (kL a), which is described as eq
kL a =
eq
non-eq
ln(CO2(air) − CO2(in) /CO2(air) − CO2(eff ) ) t
,
(1) eq
where kL a = mass transfer coefficient, 1/time, CO2(air) = CO∗2 concentration of water in equilibrium with the air entering the air-lift, CO2(in) = CO∗2 concentration of influent water, non-eq ∗ CO2(eff ) = CO2 concentration of effluent water prior to kinetic equilibration reactions, t = time. The second measure of stripping efficiency was the CO2 stripping efficiency (% CO2 stripped in a single pass), which is defined according to Grace and Piedrahita (1993, p. 501) as CO2 stripping efficency (%) =
CO2 − CO2 × 100, CO2 − CO2
(2)
eq
CO2(eff ) = CO∗2 where CO2(in) = influent CO∗2 concentration, concentration of effluent water after equilibration reactions, eq CO2(eff ) = CO∗2 concentration of water in equilibrium with the injected air. Carbon dioxide stripping efficiency differs conceptually from kL a in that uses the CO∗2 concentration of the effluent water after equieq librium reactions (CO2(eff ) ) to calculate how effectively degassing proceeded. Carbon dioxide stripping efficiency therefore combines the mass transfer process and chemical kinetic equilibria. The methodology used to derive kL a and CO2 stripping efficiencies from the recorded CO2 concentration data is described in detail in Moran (2010), with some minor differences in kL a calculations due to the nature of the degassing methodology (i.e. packing height is used to standardise kL a data for cascade columns, whereas time is used to standardise air-lift data). A general description of the calcueq lation of kL a is as follows. First, a plot was made of CO2(air) − CO2(in) eq
non-eq
versus CO2(air) − CO2(eff ) (i.e. the numerator versus denominator of the log deficit ratio term of Eq. (1)), and the slope of the regression used to calculate the log deficit ratio (R2 values for regressions eq >0.98). The calculation of CO2(air) requires knowledge of the CO2 concentration of air being injected into the air-lift. In the current study a value of 380 ppm CO2 was used based on current atmospheric observations from the NOAA Mauna Loa Observatory, Hawaii. The solubility constants for CO2 in fresh and saline water
124
D. Moran / Aquacultural Engineering 43 (2010) 120–127
were calculated from the equations given by Weiss (1974). To complete the calculation of kL a the residence time of water within the air-lift was needed to account for the time component of Eq. (1) (Reinemann and Timmons, 1989). The volume of the air-lift was calculated for the different lift heights (2.21, 2.29 and 2.44 L at 11, 16, and 25 cm lift height, respectively), allowing the residence time to be derived by dividing air-lift volume (L) by Qw (L s−1 ). The residence time (s) was used to calculate kL a, with units of s−1 . Finally, kL a was converted to a standard reference temperature of 20 ◦ C ((kL a)20 ) using the Arrhenius temperature relation kL a = (kL a)20 Temp−20
(3)
where is the temperature correction factor (=1.024), and Temp is the water temperature (◦ C) (Sincero and Sincero, 2003). The provision of kL a data allows readers to calculate the predicted airlift CO2 stripping efficiency for waters of different influent CO2 concentrations, salinities, temperatures and alkalinities using the carbonate chemistry approach outlined in Moran (2010). Unlike kL a, CO2 stripping efficiency varies according to influent CO∗2 concentration, therefore, CO2 stripping efficiency was calculated at two reference CO2(in) concentrations, 10 and 40 mg L−1 . As CO2 stripping efficiency can only be derived from the CO∗2 conceneq tration of the effluent water after equilibrium reactions (CO2(eff ) , which was not measured), this study used the same approach of the companion paper (Moran, 2010) to inter-convert CO∗2 concentration and total inorganic carbon concentration at equilibrium. The principle was to: (a) calculate the influent total inorganic carbon (Ct(inf) ) concentration based on the CO∗2 concentration and alkalinity; (b) calculate the mass of CO2 removed during degassing using the kL a mass transfer data; (c) derive the effluent total inorganic eq carbon (Ct(eff) ) concentration; (d) calculate CO2(eff ) from Ct(eff) using knowledge of the carbonate equilibrium relationship; and (e) use Eq. (2) to calculate the CO2 stripping efficiency. To complete part (a) Eqs. (13) and (14) of Moran (2010) were used to derive Ct(inf) from CO2(in) . Note that the water temperature, alkalinities and salinities were the same in both studies, thereby allowing for the use of the Ct -CO∗2 conversion functions given in Moran (2010). In part (b) the non-eq expected value of CO2(eff ) was calculated for a given value of CO2(in) eq
using Eq. (1), as the variables kL a, CO2(air) , CO2(in) , and t (residence time) were known for each salinity, lift height and Qw combination. non-eq The mass of CO2 degassed could be obtained via CO2(in) − CO2(eff ) . The results of parts (a) and (b) could then be used to calculate part (c), namely Ct(eff) . In part (d), Eqs. (16) and (17) of Moran (2010) were used to obtain an equivalent CO∗2 concentration at equilibrium eq for Ct(eff) . Finally, in part (e), the CO2(eff ) values from part (d) could eq
be used along with CO2(in) and CO2(air) to calculate CO2 stripping efficiency using Eq. (2). 2.4. Analysis of air-lift data The airflow recorded from the rotameter was corrected to standard temperature pressure using the ideal gas law, with an assumed compressibility factor of Z = 1 given the relatively low pressures used. The data for Qw and Qg were then plotted on a graph along with the logarithmic curves described by Loyless and Malone (1998) for other lift heights. The relationship between Qw and Qg in the present study was not suitably described by the logarithmic function of Loyless and Malone (1998), so a number of other functions were tested using the curve fitting software TableCurve (v4, Jandel Scientific Software, CA, U.S.A.). The only functions with a high correlation coefficient were too over-parameterised to be useful, so no function was fitted. The gas:liquid ratio (G:L) was plotted against Qw to help interpret the importance of relative gas volume in mass transfer. The pumping efficiency was evaluated using the
Fig. 4. Relationship between log deficit ratio (LDR) and three key airlift variables: lift height, pumping efficiency (G:L ratio) and water discharge rate (Qw ). (a) Dependence of LDR on lift height and G:L ratio. LDR = 0.018 + 0.010 lift + 0.111 G:L, R2 = 0.87. (b) Dependence of LDR on lift height and Qw LDR = 0.462 − 0.014 lift − 0.009 Qw + 7.23 × 10−4 lift2 + 7.82 × 10−5 Qw 2 , R2 = 0.85.
Nicklin (1963) efficiency equation cited in Reinemann et al. (1990, Eq. 14), which is valid to within 1% for air-lifts with less than 5 m submergence and a diameter greater than 20 mm. Air-lift pumping efficiency (, dimensionless) is defined as the net work done lifting the liquid divided by the work the air does as it expands isothermally Nicklin (1963). Both the log deficit ratio and the (kL a)20 were used as measures of the mass transfer performance of the air-lift. The log deficit ratio represents the proportion of CO2 transferred in the time it takes for water to travel from the inlet to the outlet of the air-lift. The (kL a)20 is the log deficit ratio divided by the residence time of water within the air-lift. The former measure of mass transfer is useful for understanding the general efficiency of an air-lift, while the latter is a standard unit used in the engineering literature and cross-study comparisons. The mass transfer performance results were plotted against Qw for each lift height, and an appropriate trend curve fitted to the combined salinity data set. The CO2 stripping efficiency data were plotted at two reference influent CO∗2 concentrations, 10 and 40 mg L−1 . The CO2 degassing data from the current study was converted to a format that could be compared to the Loyless and Malone (1998) data, which to the author’s knowledge is the only other published study on the CO2 stripping efficiency of an air-lift. Loyless and Malone (1998) published CO2 degassing data for a lift height of 15.24 cm. The (kL a)20 values were not given by Loyless and Malone (1998), but rather the degassing efficiency was expressed as the standard CO2 transfer rate (SCTR, g CO2 h−1 ). SCTR = (kL a)20 [Cm − (Cs )20 ]V 60 × 10−3
(4)
D. Moran / Aquacultural Engineering 43 (2010) 120–127
125
Fig. 5. Comparison of CO2 stripping efficiency versus Qw at different salinities and lift heights. The stripping efficiency at 15 ◦ C is given for two influent CO2 concentrations, 40 and 10 mg L−1 .The lift heights are shown at the top centre of each graph.
where (Cs )20 is saturation concentration at 20 ◦ C (0.5 mg L−1 ), Cm is standard measured concentration (arbitrarily defined as 1 mg L−1 ), V is volume of water (L), 60 is conversion from min to h, and 10−3 converts from mg to g. The SCTR was calculated for the air-lift data of the current study using the same values of (Cs )20 and Cm in order to allow for direct comparison. The SCTR data of Loyless and Malone (1998) was expressed as a function of Qw and plotted alongside the results from the current study.
3. Results and discussion The rate of water discharge from the air-lift increased in a nonlinear manner with increasing airflow (Fig. 2a), though the function that best explained the relationship was indeterminable. The relationship between Qg and Qw observed in the present study was different to that reported by Loyless and Malone (1998), whose data showed a clear logarithmic correlation for an air-lift of almost identical proportions. In the current study, there was a rapid increase in Qw from the lowest gas flow rate of 20 L min−1 , followed by a more curvilinear increase in Qw after Qg reached 30 L min−1 (Fig. 2a). It appeared that 20–30 L min−1 represented a threshold Qg that was required to initiate water flow past the level of the horizontal discharge section. This was possibly an artefact of the flowmeter used in the present study, which was only specified for a range of 20–200 L min−1 . Loyless and Malone (1998) used three airflow meters with differing operating ranges, so were probably able to obtain more accurate estimates of Qg at low airflow rates (i.e. <20 L min−1 ).
Increasing lift height reduced Qw , but there was disagreement with the relative positioning of the lift height curves between the present study and that of Loyless and Malone (1998) (Fig. 2a). In the present study, the values of Qw at each lift height tended to be higher for a given value of Qg above 30 L min−1 than that given by Loyless and Malone (1998). The salinity of the water used for the pumping efficiency measurements in the current study (35‰ NaCl) differed from the freshwater used by Loyless and Malone (1998), and the lack of comparative studies make it difficult to precisely determine the influence of salinity on air-lift pumping performance. There was consensus between the two studies that at lift heights above 11 cm it appears impossible to achieve water flow rates above c.110 L min−1 for the specified air-lift. There was a positive parabolic relationship between G:L ratio versus Qw (Fig. 2b), and a negative parabolic relationship between pumping efficiency and Qw (Fig. 2c). The G:L ratio was high (∼1.9–2.0) at low Qw and represented the initial input energy required to raise the water up the vertical riser section to the discharge pipe. The air-lift increased in pumping efficiency rapidly thereafter, and was at maximum efficiency at 50–70 L min−1 water flow (Fig. 2c). After this point the G:L ratio increased with Qw and the pumping efficiency decreased. The pumping efficiency at the minimum and maximum water flow rates were 25–40% lower than the maximum recorded efficiencies (Fig. 2c). Salinity had no obvious or consistent effect on the CO2 mass transfer characteristics of the air-lift (i.e. log deficit ratio and (kL a)20 ), with data points distributed evenly about the combined data curves (Fig. 3). There was some evidence of a higher mass transfer efficiency in freshwater at high aeration rates (i.e.
126
D. Moran / Aquacultural Engineering 43 (2010) 120–127
>100 L min−1 Qw , Fig. 3), however, there were not enough data points in this flow range to objectively test this trend. Ruen-ngam et al. (2008) investigated the effect of salinity on bubble size and oxygen mass transfer in an air-lift contactor (bubble column), and found that increasing seawater salinity resulted in smaller bubble diameters and a higher specific surface area for mass transfer. This finding intuitively suggests that mass transfer should proceed more efficiently in saltwater compared to freshwater. However, Ruen-ngam et al. (2008) only found differences in (kL a)20 between salinities at the highest aeration rates, and at the highest aeration rates mass transfer was greater in freshwater. Further testing of the air-lift in the current study at high aeration rates may yield similar salinity effects on mass transfer as that reported by Ruen-ngam et al. (2008). The concentration driving force varied with Qw in a parabolic manner, such that it initially decreased with Qw to a minimum at approximately 50 L min−1 , before increasing again (Fig. 3). The concentration driving force at the low and high end of the Qw range was approximately 75–100% greater than that at the medium Qw range (40–60 L min−1 ). This general relationship was the same for all lift heights and salinities. The parabolic relationship between log deficit ratio and Qw can be traced to the G:L ratio the air-lift was operating at. When the G:L ratio was increased, there was more air present per unit water to equalize the CO2 concentration gradient, thereby leading to greater mass transfer. The dependence of CO2 mass transfer on the G:L ratio was evident in Fig. 4a, as was the effect of lift height on CO2 mass transfer. Increasing lift height resulted in increased CO2 mass transfer at all water flow rates. The likely explanation for this is that the total contact time between the air and water increases with lift height, thereby allowing for more complete CO2 mass transfer across the concentration gradient. Comparison of the coefficients that describe log deficit ratio as a function of lift height and G:L ratio (Fig. 4a) indicated that a 5 cm increase in lift height was approximately equal to G:L increase of 0.5. Fig. 4b shows the relationship between Qw , lift height and log deficit ratio should a reader wish to calculate the CO2 mass transfer coefficient for a given Qw and lift height. The (kL a)20 values increased exponentially with Qw (Fig. 3), implying that mass transfer proceeded more rapidly as the volume of injected gas increased. In accordance with the log deficit ratio results, the CO2 stripping efficiency changed with Qw in a parabolic manner, and increasing lift height increased CO2 stripping efficiency (Fig. 5). Salinity had little effect on CO2 stripping efficiency, however, CO2 stripping efficiency was slightly higher at elevated influent CO2 concentrations. At an influent concentration of 40 mg L−1 the CO2 stripping efficiency was 1–3% higher than at an influent of 10 mg L−1 (Fig. 5). The relatively minimal effect of water salinity and influent CO∗2 concentration on the CO2 stripping efficiency of the air-lift stands in contrast to the significant effects these variables had on the stripping efficiency of a cascade column in a companion study (Moran, 2010). The reason for the discrepancy in the effects of salinity and influent CO∗2 concentration lay in the magnitude of the mass transfer efficiencies recorded in each study. The log deficit ratio values recorded for the cascade column tested by (Moran, 2010) were in the range of 1.8–2.3, which were much higher than the values measured for the air-lift in the current study (0.18–0.50, Fig. 3). The higher log deficit ratio of the cascade column represents increased mass transfer of CO2 , which considerably decreases the Ct and H+ concentration of the effluent water. The sizeable decrease in Ct and H+ concentration means that the effluent water has markedly different ionization fractions of inorganic carbon species compared to the influent water. Salinity-dependent differences in the ionization fractions of inorganic carbon species are greatest when the Ct concentration is low (relative to the alkalinity, Fig. 1 of Moran, 2010). The result is that for any given influent CO∗2 concentration,
Fig. 6. Comparison of standard CO2 transfer rates (SCTR) given by Loyless and Malone (1998) and the current study. The air-lift of Loyless and Malone (1998) was set to a lift height of 15.2 cm and was nearly identical in construction to that used in the current study. The function used to represent the data of Loyless and Malone (1998) is SCTR = 0.25 + 0.007 exp((Qw + 89)/40).
waters of different salinities exit a highly efficient degasser with different post-equilibrium effluent CO∗2 concentrations. However, in a less efficient degasser there is only a moderate decrease in Ct and H+ concentration, therefore, the ionization fractions of the influent and effluent water only differ a little, and influent CO∗2 concentration and salinity have less of an effect on CO2 stripping efficiency. The standard CO2 transfer rates increased exponentially with Qw (Fig. 6). The SCTR trend curves for lift heights 11 and 16 cm were similar, while at 25 cm lift height there was an increased SCTR for any given value of Qw (Fig. 6). The SCTR data from Loyless and Malone (1998) for a nearly identical air-lift at 15 cm lift height did not correlate closely with the data of the current study at lowmedium values of Qw . The Loyless and Malone (1998) trend line gave substantially higher SCTR values for water flow rates below 90 L min−1 (Fig. 6). The reason for the discrepancy between the studies is unclear, as Loyless and Malone (1998) did not report (kL a)20 values (upon which SCTR is reliant), nor the method used to calculate the mass of CO2 degassed.
D. Moran / Aquacultural Engineering 43 (2010) 120–127
127
4. Conclusions
References
There were significant differences in the pumping and CO2 degassing characteristics of the air-lift in the present paper compared to that of Loyless and Malone (1998), despite both studies using air-lifts of almost identical construction. The reasons for the differences are unclear and warrant further investigation for the purpose of consensus building. Salinity was not found to have a significant effect on the CO2 mass transfer characteristics of the airlift tested. In contrast to a companion study of a cascade column, the CO2 stripping efficiency of the air-lift did not differ with salinity, which was attributed to the comparatively poor mass transfer performance of the air-lift. The pumping efficiency of the air-lift was highest when Qw was between 60 and 70 L min−1 , whereas the absolute rate of CO2 transfer increased exponentially with Qw . A general equation was derived for the airlift used in this study that allows one to calculate the mass transfer coefficient for a given lift height, Qw , or G:L ratio. The mass transfer coefficient can then be used to calculate the CO2 stripping efficiency for any water type (i.e. temperature, alkalinity, salinity and influent CO2 concentration).
Grace, G.R., Piedrahita, R.H., 1993. Carbon dioxide control with a packed column aerator. In: Wang, J.K. (Ed.), Techniques for Modern Aquaculture. American Society of Agricultural Engineers, Saint Joseph, MI, pp. 496–505. Grace, G.R., Piedrahita, R.H., 1994. Carbon dioxide control. In: Timmons, M.B., Losordo, T.M. (Eds.), Aquaculture Water Reuse Systems: Engineering Design and Management. Elsevier, New York, USA, pp. 209–234. La Motta, E.J., 1995. Chemical analysis of CO2 removal in tray aerators. J. Am. Water Resour. Assoc. 31, 207–216. Loyless, J.C., Malone, R.F., 1998. Evaluation of air-lift pump capabilities for water delivery, aeration, and degasification for application to recirculating aquaculture systems. Aquacult. Eng. 18, 117–133. Moran, D., 2010. Carbon dioxide degassing in fresh and saline water. I. Degassing performance of a cascade column. Aquacult. Eng. 43, 29–36. Moran, D., Tirsgård, B., Steffensen, J.F., 2010. The accuracy and limitations of a new meter used to measure aqueous carbon dioxide concentration. Aquacult. Eng. 43, 101–107. Nicklin, D.J., 1963. The air-lift pump: theory and optimization. Trans. Inst. Chem. Eng.. Reinemann, D.J., Timmons, M.B., 1989. Prediction of oxygen transfer and total dissolved gas pressure in airlift pumping. Aquacult. Eng. 8, 29–46. Reinemann, D.J., Parlange, J.Y., Timmons, M.B., 1990. Theory of small-diameter airlift pumps. Int. J. Multiphase Flow 16, 113–122. Ruen-ngam, D., Wongsuchoto, P., Limpanuphap, A., Charinpanitkul, T., Pavasant, P., 2008. Influence of salinity on bubble size distribution and gas–liquid mass transfer in airlift contactors. Chem. Eng. J. 141, 222–232. Sincero, A.P., Sincero, G.A., 2003. Chapter 9: Aeration, Absorption and Stripping, Physical–Chemical Treatment of Water and Wastewater. CRC Press, Florida, USA. Stumm, W., Morgan, J.J., 1996. Aquatic Chemistry: Chemical Equilibria and Rates in Natural Waters. Wiley-Interscience, New York, USA. Summerfelt, S.T., Vinci, B.J., Piedrahita, R.H., 2000. Oxygenation and carbon dioxide control in water reuse systems. Aquacult. Eng. 22, 87–108. Summerfelt, S.T., Davidson, J., Waldrop, T., 2003. Evaluation of full-scale carbon dioxide stripping columns in a coldwater recirculating system. Aquacult. Eng. 28, 155–169. Timmons, M.B., Ebeling, J.M., Wheaton, F.W., Summerfelt, S.T., Vinci, B.J., 2002. Recirculating Aquaculture Systems. Cayuga Aquaculture Ventures, Ithaca, NY, 769 pp. Watten, B.J., Sibrell, P.L., Montgomery, G.A., Tsukuda, S.M., 2004. Modification of pure oxygen absorption equipment for concurrent stripping of carbon dioxide. Aquacult. Eng. 32, 183–208. Weiss, R.F., 1974. Carbon dioxide in water and seawater: the solubility of a non-ideal gas. Mar. Chem. 2, 203–215.
Acknowledgements I am thankful for the technical and practical support given by M. Fülberth, and logistical support given by J. Bregnballe, H. Jarlbæk, B.H Olsen, and J.G. Støttrup. I also thank the three anonymous reviewers for reviewing and improving this manuscript. This work was partially funded by the EU Fisheries Sector Program FIUF and the Danish Ministry of Food, Agriculture and Fisheries. The author was supported by a fellowship from the New Zealand Foundation for Research Science and Technology.