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Gas-lift circulation of a liquid between two inter-connected bubble columns
Journal Pre-proofs Gas-lift circulation of a liquid between two inter-connected bubble columns Mehdi Jafarian, Yusuf Chisti, Graham J. Nathan PII: DOI: Reference:
S0009-2509(20)30106-8 https://doi.org/10.1016/j.ces.2020.115574 CES 115574
To appear in:
Chemical Engineering Science
Received Date: Revised Date: Accepted Date:
25 October 2019 7 February 2020 11 February 2020
Please cite this article as: M. Jafarian, Y. Chisti, G.J. Nathan, Gas-lift circulation of a liquid between two interconnected bubble columns, Chemical Engineering Science (2020), doi: https://doi.org/10.1016/j.ces.2020.115574
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© 2020 Published by Elsevier Ltd.
Gas-lift circulation of a liquid between two inter-connected bubble columns
Mehdi Jafarian a,*, Yusuf Chisti b, Graham J. Nathan a
a
Centre for Energy Technology, School of Mechanical Engineering, The University of
Adelaide, SA 5005, Australia
b
School of Engineering, Massey University, Private Bag 11 222, Palmerston North, New
Zealand
* Corresponding author: School of Mechanical Engineering, The University of Adelaide, Adelaide, SA 5005, Australia. Tel.: +61-8-8303-5460; E-mail: [email protected] (M. Jafarian).
1
Abstract A pneumatically-mixed novel configuration of two inter-connected bubble column reactors is reported using air-water to demonstrate the device, enabling circulation of a liquid between the columns solely by a gas-lift mechanism. The new reactor may be used for circulating a high temperature/corrosive liquid between the columns, enabling the cyclical processing of a liquid medium with two different gaseous reactants. The effects of superficial gas velocity, unaerated liquid height in the columns and the diameter of aeration nozzles are discussed with reference to the gas holdup in, and rate of liquid circulation between, the two columns, together with the energetics of circulation. The height of the gas-free liquid in the reactor was found to influence the liquid circulation velocity positively, but the effect of aeration nozzles diameter was marginal. The lowest and highest measured flow rates of circulating water between the columns were 0.7 dm3 min1 and 4.2 dm3 min1, respectively.
Keywords: airlift reactor; bubble column; chemical looping; two-phase flow; gas holdup.
2
Nomenclature A
cross sectional area of a column (m2)
AC1
cross sectional area of column 1 (m2)
AC2
cross sectional area of column 2 (m2)
a
parameter in Eq. (20) (dm3 min1)
aL
gas–liquid interfacial area per unit volume of liquid (m1)
b
parameter in Eq. (20) (-)
d
internal diameter (ID) of gas injection nozzle (m)
EF
frictional energy dissipation rate in the entire circulation loop (W)
EK,C1 kinetic energy input rate in column 1 (W) EK,C2 kinetic energy input rate in column 2 (W) ES
energy dissipation rate due to stagnant gas in downflow ducts (W)
EW,C1 energy dissipation rate in the liquid wakes behind the bubbles in column 1 (W) EW,C2 energy dissipation rate in the liquid wakes behind the bubbles in column 2 (W) EX,C1
energy input rate due to isothermal gas expansion in column 1 (W)
EX,C2
energy input rate due to isothermal gas expansion in column 2 (W)
g
gravitational acceleration (m s2)
HL
height of gas-free liquid in a bubble column (m)
Hm1
manometer height shown in Fig. 1 (m)
Hm2
manometer height shown in Fig. 1 (m)
ΔHm
manometer reading (m)
hD
height of gas-liquid dispersion in a column (m)
hD,C1
height of gas-liquid dispersion in column 1 (m)
hD,C2
height of gas-liquid dispersion in column 2 (m)
mG
mass flow rate of gas in a column (kg s1) 3
Ph
headspace pressure (Pa)
Qc
flow rate of the circulating liquid between bubble columns (dm3 min1)
Qm
molar flow rate of air in the column (kmol s1)
R
gas constant (kJ kmol1 K1)
T
absolute temperature (K)
UL
superficial liquid velocity in bubble column (m s1)
UL,C1 superficial liquid velocity in column 1 (m s1) UL,C2 superficial liquid velocity in column 2 (m s1) Us
superficial air velocity in bubble column (m s1)
Us,C1
superficial air velocity in column 1 (m s1)
Us,C2
superficial air velocity in column 2 (m s1)
un
gas velocity at the injection nozzle (m s1)
VG
volume of the gas phase (m3)
VL
volume of the liquid phase (m3)
VL,C1
volume of liquid in column 1 (m3)
VL,C2
volume of liquid in column 2 (m3)
Vt
total volume of gas-liquid dispersion in an aerated column (m3)
Greek letters
G
gas holdup (-)
G,C1
gas holdup in column 1 (-)
G,C2
gas holdup in column 2 (-)
contribution of gas expansion to total power input (%)
ρL
density of the liquid (kg m3)
4
ρG
density of gas (kg m3)
fraction of jet’s momentum energy transferred to liquid (-)
r
mean residence time of the liquid in a bubble column (s)
5
1. Introduction Gas-lift reactors are a well-established technology in gas-liquid contacting operations, employing the action of a compressed gas to achieve good mixing and high rates of heat and mass transfer in a liquid or slurry (Merchuk and Siegel, 1988; Chisti, 1998; Merchuk, 2003; Zhang et al., 2017). The avoidance of moving parts, such as shafts and stirrers, reduces capital and operational costs (Chisti and Moo-Young, 1987; Jafarian et al., 2019) and, in biological processes, reduces the potential for contaminating microorganisms entering the vessel through shaft seals. Conventional gas-lift reactors, such as those used in biological treatment of wastewater and other bioprocessing applications, typically contact only a single gas (e.g. air) with the liquid, or slurry, being mixed. However, much less attention has been paid to bubbling reactors that enable the looping of reactants between two vessels. Various geometric configurations of multi-column gas-lift reactors have been described previously (Chisti and Moo-Young, 1987; Chisti, 1998; Merchuk, 2003). In all cases, the reactor consists of two relatively tall vertical zones that are linked near to the top and bottom. One of these vertical zones (the riser) is typically bubbled with a gas. This results in an up flow of the gas-liquid dispersion in the riser zone and a downflow in the interconnected vertical downcomer zone that is mostly free of gas (Merchuk and Siegel, 1988; Merchuk, 2003). This circulatory flow is driven mostly by the difference in bulk densities between the gas-liquid dispersion in the riser and that in the downcomer zones (Chisti and Moo-Young, 1987). In specific cases, the momentum of the compressed gas injected near to the bottom of the riser zone may also contribute to circulation of the fluid. In most of these gas-lift reactors, the configuration of the top zone where the riser and the downcomer connect, ensures that most of the gas disengages from the fluid before it enters the downcomer. Gas disengagement is necessary for achieving high rates of liquid circulation. Often the top zone is purpose-designed to enhance gas-liquid separation (Chisti and Moo-Young, 1993; Chisti, 1998). Thus, while this
6
system is suitable for performing a single reaction between a gas and liquid or slurry, by itself it cannot be used to perform two cyclical reactions as is needed for chemical looping processes. To address the above gap, a novel gas-lift configuration capable of achieving this has been recently developed by Jafarian et al. (2017a; 2017b). This device uses dual interconnected gas-lift reactors (Fig. 1) to circulate a liquid between two reactors, each bubbling a different gaseous reactants, to be suitable for chemical looping processes. As shown in Fig. 1, this novel configuration is comprised of two interconnected gas-sparged bubble column riser zones. The fluid leaving the top of each column passes through degassing zones to enter the other column near the bottom to produce an -shaped flow loop (Fig. 1). The degassing zones also prevent mixing between the different gaseous reactants by assuring disengagement of the bubbles from the liquid exiting a bubble column and before entry to the other column. Two different reactant gases can therefore be used simultaneously without being mixed. The absence of moving parts in the reactor (Fig. 1) is expected to improve its utility under challenging processing conditions involving high temperatures and corrosive reactants that can be difficult to circulate with mechanical pumping. Examples include molten metal oxides at high temperature, which are used as oxygen carriers in liquid chemical looping in combustion/gasification systems (Jafarian et al., 2017b; Sarafraz et al., 2017a; Sarafraz et al., 2017b; Sarafraz et al., 2018). Similarly, the circulation of molten bromide salt is used in a chemical looping process for oxidation of hydrogen bromide to bromine (Upham et al., 2017). Although airlift-driven tubular-loop photobioreactors conceptually similar to that shown in Fig. 1, have been reported previously (Lee et al., 1995), no data are presently available on their operation or hydrodynamic characteristics (Fig. 1). Therefore, the first objective of the present work was to meet this need. Although the factors that influence circulation rate of liquids in airlift reactors are well known for single riser configurations, their relative significance in gas-lift reactors with dual 7
inter-connected riser zones (Fig. 1), is not. Liquid circulation rate affects the residence time of the fluid in different zones of a reactor such as the one shown in Fig. 1 and, hence, the kinetics of any reactions being carried out (Zhang et al., 2005). For example, the degree of oxidation of fuel in chemical looping combustion and gasification systems has been shown to depend on the recirculation flow rate of the molten oxygen carrier between the fuel and air compartments of the reactor (Jafarian et al., 2017b; Sarafraz et al., 2017a; Sarafraz et al., 2017b; Sarafraz et al., 2018; Sarafraz et al., 2019a; Sarafraz et al., 2019b). On this basis, the aim of the present investigation is to advance the understanding of the factors that control the circulation rate in a dual interconnected bubble column system and to provide an experimental assessment of the hydrodynamics and liquid circulation characteristics of the reactor configuration shown in Fig. 1.
2. Methodology 2.1. Theoretical analysis All energy for circulation of the liquid in the two-column -loop reactor system (Fig. 1) is generated by the gas injected into the two bubble columns. Part of the energy input is a consequence of the work done during isothermal expansion of the gas as it moves up the columns, while the remainder is a consequence of the kinetic energy of the gas jets issuing from the sparger nozzles. Energy is dissipated in the wakes of the circulating liquid behind the bubbles (Chisti et al., 1988; Chisti, 1989) that rise in the bubble columns and from the changes in cross section of the flow path and friction in the flow circuit including the connecting pipes and the gas disengaging devices (Fig. 1). An energy balance over the circulation loop (Fig. 1) can be written as follows: 𝐸𝑋,𝐶1 + 𝐸𝑋,𝐶2 + 𝐸𝐾,𝐶1 + 𝐸𝐾,𝐶2 = 𝐸𝑊,𝐶1 + 𝐸𝑊,𝐶2 + 𝐸𝐹 + 𝐸𝑆
8
(1)
In the above equation: EX,C1 and EX,C2 are the energy input rates due to isothermal expansion of the gas in columns 1 and 2, respectively; EK,C1 and EK,C2 are the kinetic energy input rates due to gas injected at the bottoms of columns 1 and 2, respectively; EW,C1 and EW,C2 are energy dissipation rates in the liquid wakes behind the bubbles rising in columns 1 and 2, respectively; EF is the energy dissipation rate due to cross sectional changes and friction in the circulation loop; and ES is the energy dissipation rate due to any stagnant gas (Chisti et al., 1988; Chisti, 1989) in the two downflow ducts connecting the bubble columns (Fig. 1). Equation (1) disregards the kinetic energy dissipation rate at the top openings of the bubble columns because of the low gas flow velocity in those regions. It also ignores frictional losses in the bubble columns because of the low liquid flow rates in them. The ES term in Eq. (1) can be neglected as no stagnant gas bubbles were observed in the interconnecting ducts because of the disengagement of gas in the gas-liquid separators (Fig. 1). The energy dissipation rate in the wakes behind the bubbles in column 1 (EW,C1) and column 2 (EW,C2) of the circulation loop was calculated using the following equations (Chisti et al., 1988): 𝐸𝑊,𝐶1 = 𝐸𝑋,𝐶1 + 𝐸𝐾,𝐶1 ― 𝜌𝐿𝑔𝑈𝐿,𝐶1𝜀𝐺,𝐶1ℎ𝐷,𝐶1𝐴𝐶1
(2)
𝐸𝑊,𝐶2 = 𝐸𝑋,𝐶2 + 𝐸𝐾,𝐶2 ― 𝜌𝐿𝑔𝑈𝐿,𝐶2𝜀𝐺,𝐶2ℎ𝐷,𝐶2𝐴𝐶2
(3)
where L is the density of the liquid, g is gravitational acceleration, UL is the superficial liquid velocity, G is the gas holdup, hD is the height of gas-liquid dispersion in the column, and A is the cross sectional area of the column. The subscripts C1 and C2 denote columns 1 and 2, respectively. The total volume Vt of gas-liquid dispersion in an aerated column can be related to the height of dispersion (hD) in the column and the cross sectional area (A), as follows: (4)
𝑉𝑡 = ℎ𝐷𝐴 where Vt is the sum of the volumes of the gas (VG) and liquid phases (VL), or: 9
(5)
𝑉𝑡 = 𝑉𝐺 + 𝑉𝐿
Furthermore, by definition, the gas holdup (G), or volume fraction of gas in gas-liquid dispersion, depends on the volumes of the individual phases, as follows: 𝜀𝐺 =
𝑉𝐺
(6)
𝑉𝐺 + 𝑉𝐿
From Eq. (5) and Eq. (6), the following equation can be obtained:
(
𝑉𝑡 = 𝑉𝐿
)
1 1 ― 𝜀𝐺
(7)
Substitution of Eq. (4) and Eq. (7) into Eq. (2), leads to the following: 𝐸𝑊,𝐶1 = 𝐸𝑋,𝐶1 + 𝐸𝐾,𝐶1 ― 𝜌𝐿𝑔𝑈𝐿,𝐶1𝑉𝐿,𝐶1
(
𝜀𝐺,𝐶1
)
(8)
)
(9)
1 ― 𝜀𝐺,𝐶1
Analogously, for column 2 we obtained: 𝐸𝑊,𝐶2 = 𝐸𝑋,𝐶2 + 𝐸𝐾,𝐶2 ― 𝜌𝐿𝑔𝑈𝐿,𝐶2𝑉𝐿,𝐶2
(
𝜀𝐺,𝐶2
1 ― 𝜀𝐺,𝐶2
With ES = 0, as discussed above, the substitution of Eq. (8) and Eq. (9) into Eq. (1), provides the following expression for energy dissipation rate (EF) in the loop:
[
𝐸𝐹 = 𝜌𝐿𝑔 𝑈𝐿,𝐶1𝑉𝐿,𝐶1
(
𝜀𝐺,𝐶1
)+𝑈
1 ― 𝜀𝐺,𝐶1
)]
(
𝜀𝐺,𝐶2 𝐿,𝐶2𝑉𝐿,𝐶2 1 ― 𝜀𝐺,𝐶2
(10)
The percentage () of the total input power that is transferred to gas expansion were calculated as follows: 𝐸𝑋,𝐶1 + 𝐸𝑋,𝐶2
(11)
𝜃 = 𝐸𝑋,𝐶1 + 𝐸𝑋,𝐶2 + 𝐸𝐾,𝐶1 + 𝐸𝐾,𝐶2 × 100
In the above equation, the power input resulting from gas expansion (i.e. EX,C1, EX,C2) and the kinetic power input (i.e. EK,C1, EK,C2), were calculated using the following equations (Chisti, 1989): 𝐸𝑋 = 𝜌𝐿𝑔𝑈𝑠𝑉𝐿
(12)
𝐸𝐾 = 0.5 𝛺 𝑚𝐺𝑢2𝑛
(13)
where 𝑚𝐺 is the mass flow rate of gas in a column, un is the gas velocity at the injection nozzle and is the fraction of the jet’s kinetic energy that is transferred to the liquid. Based on earlier 10
work, is estimated to be approximately 0.06 (Moo-Young and Blanch, 1981; Chisti, 1989). The superficial gas velocity in the above equations was calculated using the following equation (Chisti, 1989): 𝑄𝑚𝑅𝑇
(
𝑈𝑠 = 𝑉𝐿𝜌𝐿𝑔𝑙𝑛 1 +
)
𝜌𝐿𝑔𝐻𝐿 𝑃ℎ
(14)
where Qm is the molar flow rate of air, R is the gas constant, T is the absolute temperature, VL is the volume of liquid in a column, and Ph is the headspace pressure in the columns open to atmosphere. The mean residence time (r) of the liquid in each bubble column, was calculated using the following equation (Merchuk, 2003; Zhang et al., 2017): 𝜏𝑟 =
𝑉𝐿
(15)
𝑄𝑐
where Qc is the flow rate of the circulating liquid between the bubble columns. The reactor shown in Fig. 1 can potentially be operated in two alternative modes: either (1) with equal values of superficial air velocities in the two columns, or (2) with different values of superficial gas velocity in the two columns. Mode 2 may be useful if different gaseous reactants are to be used, for example (Jafarian et al., 2017b; Sarafraz et al., 2017a; Sarafraz et al., 2017b; Sarafraz et al., 2018). Both of these operating modes were investigated, as is discussed in the following sections.
2.2. Experimental reactor and measurements The reactor comprised of two identical cylindrical bubble columns (diameter = 85 mm) connected to each other as shown in Fig. 1. Liquid circulation and mixing in the reactor were studied using a deionized water–air system. Air from a compressor (Fig. 1) was sparged at the bottom of the bubble columns using identical nozzles of specified diameter. Two electronic
11
mass flow controllers were used to control the flow rate of air to the bubble columns at specified values. The air flow rate was reproducible to within 2% of the set value. Two ultrasonic flow meters (accurate to 5% of the measured value in the range of 0.2– 0.5 dm3 min1 and 3% of the measured value in the range of 0.5–25 dm3 min1) were used in parallel to measure the flow rate of the liquid flowing between the columns (Fig. 1; only one flow meter is shown for clarity). The steady-state flow was measured at 0.1 s intervals for 60 s and the data were averaged. LabVIEW software (National Instruments, Austin, TX, USA) was used to post-process the data. Gas holdup (G), or volume fraction of gas in a gas-liquid dispersion, is an important operational characteristic of gas-lift reactors for several reasons: (1) a gas-liquid reactor must be sized to accommodate the expected gas holdup (Chisti, 1989); (2) differences in gas holdup in different zones of a gas-lift reactor determine the rate of liquid circulation in the reactor loop (Chisti et al., 1988; Chisti, 1989); (3) absorption, or mass transfer, of a gaseous reactant is strongly influenced by the gas–liquid interfacial area (aL) per unit volume of liquid in the reactor (Chisti, 1989) so that aL depends on both the gas holdup and the average bubble diameter; and (4) the average residence time of gas in liquid depends on gas holdup (MooYoung and Blanch, 1988; Chisti, 1989). The gas holdup (G) was measured in both bubble columns using the manometeric method (Chisti, 1989). Two manometers were connected to the columns as shown in Fig. 1. The manometer readings (ΔHm) were recorded digitally with a Sony ILCE-6000 camera. These images were then post processed to obtain measurements of ΔHm. The ΔHm value was reproducible to within 1% of the measured value near the upper limit and to within 12% at the low end of the measured values. The average steady-state gas holdup in the column between the taps of the manometer was calculated using the following equation (Chisti, 1989):
12
𝜀𝐺 =
(
𝜌𝐿
)
𝜌𝐿 + 𝜌𝐺
𝐻𝑚2 ― 𝐻𝑚1 𝐻𝑚
=
(
𝜌𝐿
)
𝛥𝐻𝑚
(16)
𝜌 𝐿 + 𝜌 𝐺 𝐻𝑚
where ρL was the density of the liquid, ρG was the density of gas, ΔHm was the manometer reading (Hm = Hm2 – Hm1; Fig. 1) and Hm was the distance between the taps of the manometer. The experimental variables were the superficial gas velocity (Us) in the bubble columns, the height (HL) of the gas-free liquid in each column and the diameter (d) of the gas injection nozzles. The unaerated liquid height HL was identical in the two bubble columns and was level with the height of the horizontal exit port located near to the top of each column (Fig. 1). In the various experiments, HL ranged from 0.212 m to 0.600 m. Whenever HL was changed, the location of the fluid exit port in the two columns was also changed. Identical gas injection nozzles were used in the two bubble columns for each experiment, although the internal diameter of the nozzles was varied (d = 1.4, 2.0, 2.8, 4.0 and 5.6 mm). All other geometric measurements of the reactor system remained fixed as shown in Fig. 1. The specific combinations of HL, sparger nozzles and aeration velocities used in different experiments are summarized in Table 1. Experiments were carried out at room temperature (24 C) and atmospheric pressure, so that the correlations developed in this work apply to operation at atmospheric pressure.
3. Results and discussion 3.1. Operation with equal aeration velocity in the two bubble column zones Figure 2 shows that the gas holdup in the two bubble column zones depends on the superficial air velocity via a power law. This is consistent with other well-known observations about gas holdup in conventional bubble columns (Hikita et al., 1980; Chisti and Moo-Young, 1988; Chisti, 1989; Deckwer, 1992; Besagni et al., 2015; Besagni and Inzoli, 2016; Besagni and Inzoli, 2017). Similar values of g were measured in both columns (Fig. 2), which is expected because the columns were identical and operated under the same conditions. The gas 13
holdup correlations for the two columns are also identical to within measurement accuracy (Fig. 2), giving confidence in the measurements. Therefore, the two correlations were combined into the following equation: 𝜀𝐺 = 63.33 𝑈0.80 𝑠
(17)
Equation (17) was obtained by regression of the combined data for the two columns. Gas holdup showed barely any dependence on the diameter of the single-hole sparger nozzle, notwithstanding the broad range of hole diameters (d = 1.4 to 5.6 mm) tested (Fig. 2). This is consistent with observations in conventional bubble columns (Hikita et al., 1973; Hikita et al., 1980), suggesting that the gas holdup is controlled by the prevailing turbulence level in the water rather than by the initial size of the gas bubbles generated at the sparger nozzle (Chisti, 1989; Deckwer, 1992). The bubble columns in the present work (Fig. 1) had a superimposed liquid up-flow because each column also served as the riser zone of the airlift circulation loop (Fig. 1). This is unlike a conventional bubble column, which lacks a net up-flow of liquid (Chisti, 1989; Camacho Rubio et al., 2004). The measured liquid circulation rate between the columns is shown in Fig. 3. The rate of the liquid circulation and up flow of liquid in the bubble columns can be seen to increase with increasing aeration velocity (Fig. 3). An imposed cocurrent liquid flow in a bubble column reduces the gas holdup relative to the case of an otherwise identical conventional column operated without flow through of liquid (Chisti, 1989). In addition, presence of flowing liquid in a bubble column postpones, or suppresses, the transition of flow regime from the bubble-flow regime to churn-turbulent (or heterogeneous) flow regime, to higher values of the superficial gas velocity. In air–water in conventional bubble columns this flow transition occurs around a superficial gas velocity value of 0.05 m s–1 (Chisti, 1989; Chisti, 1998).
14
Figure 4 compares the gas holdup predictions of Eq. (17) with two well-known equations developed for gas holdup in conventional bubble columns operated with air–water (Chisti and Moo-Young, 1988; Chisti, 1989; Chisti, 1998). As explained above, the gas holdup in the present study was lower and the flow transition from the bubble-flow regime to churnturbulent regime did not occur at the superficial aeration velocity expected for a conventional column (Fig. 4). The two literature equations given in Fig. 4 for the bubble-flow and churnturbulent flow regimes in air–water, have correlated data from many different conventionally operated bubble columns. For the columns in the present work (Fig. 3), the liquid circulation rate was 0.7 dm3 min–1 for the lowest air flow velocity (0.01 m s1), while it was 2.8 dm3 min–1 at the highest value of aeration velocity (0.16 m s1). This range of liquid volume flow rates corresponds to a liquid superficial velocity range in a bubble column of 2.1103 to 9.6103 m s1. For fixed geometry of the circulation loop (Fig. 1) and complete disengagement of gas from the interlinking tubes, the gas holdup in the risers (i.e. the bubble columns) provides the main driving force for liquid circulation. Therefore, as the gas holdup increased with increasing aeration rate in the columns, the liquid flow velocity increased, leading to a decreasing residence time of liquid in each column (Fig. 3). Over the operational range of aeration velocities (around 0.01–0.16 m s1), the residence time varied by around 3.7-fold (from 149 to 40 s). In principle, it would be possible to have different liquid residence times in the two columns, if the columns had different diameters but were otherwise identical. As gas holdup was independent of the diameter of the aeration nozzle, barely any effect of the nozzle diameter could be seen on the induced liquid circulation rate and the liquid residence time in a column (Fig. 3). A lack of significant effect of aeration nozzle diameter on liquid circulation velocity also suggests that the kinetic energy of the air jet issuing from the sparger contributes little to liquid circulation for these conditions. 15
The energy dissipation rate (EF), calculated using the measured values of gas holdup in Eq. (10), is shown in Fig. 5 for the range of superficial air velocities in the two columns. It can be seen that EF is independent of diameter of the aeration nozzle as is the gas holdup in Fig. 2. Also EF dependence on the superficial aeration velocity is well described by an empirical power law, as follows: 𝐸𝐹 = 0.395 𝑈1.18 𝑠
(18)
This equation is consistent with Eq. (10) if the dependence of gas holdup on the superficial aeration velocity is considered. The power imparted to the liquid was contributed by isothermal expansion of the air as it rose up the bubble columns and the kinetic energy of the air jets formed at the sparger nozzles. Fig. 6 presents (calculated from Eq. 11) as a function of the superficial gas velocity. This shows that, for all but the two smallest nozzles (d 2.0 mm), 80% or more of the energy input to the liquid in the riser column is provided by isothermal expansion of the air. However, for the narrow diameter nozzles operated at high aeration rates, the contribution of the kinetic energy of the gas jet becomes significant, showing that it is possible to develop devices that operate in this regime. Nevertheless, for all cases investigated here, the contribution of gas expansion to total energy input is always at least 50%. As expected, the kinetic energy input increases as the cross sectional area of the nozzle is reduced for a given superficial gas velocity.
3.2. Operation with different aeration rates in the bubble column zones The measured liquid circulation rate (Qc) between the two bubble columns is compared in Fig. 7 for the two different modes of operation: (a) the mode with identical values of the superficial air velocities in both columns (i.e. Us,C1 = Us,C2) and (b) the mode with the value of Us,C1 fixed at 0.039 m s1 for different Us,C2 values in column 2 (Table 1). The case with Us = 0 m s1 in column 2 is equivalent to operating as a conventional airlift reactor with a single 16
aerated riser zone, despite the unusual -shaped recirculation loop. For this comparison, both column were installed with an identical sparger nozzle (d = 2.8 mm) and the height of gas-free liquid in both columns was also fixed at HL = 300 mm. The two columns have an identical energy input when the superficial aeration velocity in column 2 is 0.039 m s1 (Fig. 7). Fig. 7 reveals that, irrespective of the mode of operation, the induced liquid circulation rate depends on the total energy input to the reactor system. For Us,C2 = 0.039 m s1, the energy input rate due to aeration in both modes of operation is equal, as is the liquid circulation rate (Fig. 7). For Us,C2 > 0.039 m s1, the reactor system with both the bubble columns aerated at equal aeration velocities always has a higher total energy input rate (black line in Fig. 7) and, consequently also, a higher liquid circulation rate than the reactor loop operated with Us,C1 fixed at 0.039 m s1 (red line in Fig. 7). The situation is reversed for Us,C2 < 0.039 m s1, so that the total power input rate in reactor system (i.e. the red line in Fig. 7) is greater than the case with Us,C1 = 0.039 m s1 (i.e. the black line in Fig. 7). For all cases, the total liquid circulation flow rate increases with the total energy input. Fig. 8 presents both the liquid circulation rate (Qc) and the mean residence time of water in each column (r) as a function of the superficial aeration velocity in column 2 (Us,C2). It can be seen that a minimum rate of circulation occurs with no gas injection in column 2 (i.e. Us,C2 = 0 m s1) because of the constant aeration through column 1 (Us,c1 = 0.039 m s1). This case also corresponds to the maximum residence time through the two columns, as expected. For non-zero values of Us, both the total circulation rate and the residence time are well described by empirical correlations, so that Qc = 1.02 + 3.06Us,C20.564 and r = 105.13 113.39Us,C20.385. The values of gas holdup corresponding to the data in Fig. 8, in the two columns are presented in Fig. 9. Gas holdup in column 2 increases with increasing superficial air velocity Us,C2 in keeping with the dependence of this variable on aeration velocity discussed above. Similarly, the gas holdup remains constant at 5% in column 1, as consistent with the constant 17
aeration velocity in that column (Us,C1 = 0.039 m s1). Nonetheless, both columns contribute to the flow circulation, as discussed above. Maintaining different values of gas holdup in the two columns may be desirable for some applications, such as where two gas-liquid reactions are being carried out in these zones for chemical looping applications. The energy loss due to circulation of water for the mode of operation with unequal aeration rates in the two columns can be seen to depend on the aeration velocity (Us,C2) in column 2 as shown in Fig. 10. The EF values, as calculated using Eq. (10), are well described by the following empirical equation (Fig. 10): 𝐸𝐹 = 2.3 × 10 ―3 +0.145𝑈𝑠,𝐶2 +0.121𝑈2𝑠,𝐶2
(19)
The non-zero value of energy loss for Us,C2 = 0 m s1 (Eq. 19) can be attributed to the circulation induced by the constant aeration rate in column 1 (Us,C1 = 0.039 m s1). Relative to operation with equal superficial air velocities in the two columns (operation mode (b), Fig. 10), the mode of operation with unequal aeration has lower values of EF where Us,C2 > 0.039 m s1, but higher values otherwise. This behavior is also explained by the differences in the induced liquid flow rate in different modes of operation discussed above.
3.3. Operation with equal aeration velocity in both bubble columns: effects of the height of gas-free liquid The dependence of the circulation flow rate (Qc) between the bubble columns on superficial gas velocity for various heights of unaerated liquid in the columns is presented in Fig. 11. In any given experiment, the HL values in the two columns were identical and varied over a range of 212 to 600 mm. The measured liquid circulation flow rate (Qc) increased with the aeration velocity (Us) and HL, consistent with previous work (Chisti et al., 1988; Chisti, 1989; Chisti and Moo-Young, 1993).
18
The observations in Fig. 11 were are also consistent with those seen previously in Fig. 3 for the same mode of operation but with a specific value of HL (300 mm). The dependence of the liquid circulation rate on the aeration velocity in the two columns is well described by the following power law: 𝑄𝑐 = 𝑎 𝑈𝑏𝑠
(20)
Here, the dependence of the values of the empirical constants a and b are as follows: 𝑎 = 4.352𝑙𝑛 𝐻𝐿 +11.852
(21)
𝑏 = 0.4 𝐻𝐿―0.142
(22)
Combining these dependences, Eq. (20) can be rewritten as follows: ―0.142
𝐻𝐿 𝑄𝑐 = (4.352𝑙𝑛 𝐻𝐿 + 11.852)𝑈0.4 𝑠
(23)
This empirical equation was found to correlate all of the data in Fig. 11 except for 2 data points to within 15% of an exact agreement, as is shown in Fig. 12. The correlation is consistent with earlier studies in conventional airlift reactors, which show that the rate of induced liquid circulation increases with the height of dispersion in the reactor, for otherwise identical values of aeration rate and other geometric characteristics of the reactor (Chisti et al., 1988; Chisti, 1989). Fig. 13 presents the values of r calculated using Eq. (15) as a function of the aeration rate (identical in the two columns) for various values of HL (identical in both columns). A strong inverse relationship can be seen. However, any effect of HL on the liquid residence time is small, unlike the influence of HL on the liquid circulation rate (Fig. 11). This can be explained by the co-dependence on HL of Qc and the volume of liquid (VL in Eq. 15). Although Qc increases with increasing HL, so does VL. So that the net effect on the residence time of changing HL is minor (Fig. 13). The dependence of gas holdup on superficial velocity is presented for the two columns in Fig. 14. The response is almost identical, consistent with nominally identical values of HL 19
and aeration velocities, with the differences attributable to experimental uncertainty. The gas holdup can be seen to increase with increasing aeration velocity for a given value of unaerated liquid height in a column, as explained above. However, at any fixed aeration velocity, increasing the height of the gas-free liquid increases the gas holdup (Fig. 14). This can be possibly attributed to an increase in the liquid superficial velocity with an increase in HL (as seen in Fig. 11) which, in turn, may increase the turbulence-induced shear rate within the liquid. Such an increase in turbulent shear can be expected to reduce the mean bubble size and, therefore, also reduce the average rise velocity of the bubbles. In conventional bubble columns without an imposed liquid flow, neither the height of liquid in the column nor the column diameter influence the average gas holdup (Chisti, 1989; Sasaki et al., 2017), provided that the columns are sufficiently large (e.g. diameter 0.1 m, gasfree liquid height 1 m). In contrast, the height of gas-lift reactors profoundly affects the gas holdup and liquid circulation (Chisti et al., 1988; Chisti, 1989; Chisti and Moo-Young, 1993). In airlift reactors, the efficiency of gas-liquid separation in the top region of the reactor as well as the geometry of the degassing zone (Chisti, 1998; Chisti and Moo-Young, 1993) further impact liquid circulation rate and gas holdup (Chisti et al., 1988; Chisti, 1989; Chisti and MooYoung, 1993). The dependence on superficial gas velocity of the fraction of the total energy input () that was contributed by isothermal expansion of the gas is presented in Fig. 15 for a series of heights of gas-free liquid. While the general trend is similar to that in Fig. 6, it can be seen here that an increase in HL causes an increase in for all values of Us. This can be attributable to the increase in hydrostatic pressure at the bottom of the column with an increase in HL, meaning that the gas needs to expand from a higher pressure at the bottom of the column to reach the atmospheric pressure at the surface of the gas-liquid interface. It is also noteworthy that, with
20
the exception of a single case (the -value at HL = 212 mm and Us 0.16 m s1; Fig. 15), 80% of the total power input to the liquid is attributable to the isothermal expansion of air.
4. Conclusions A gas-lift driven circulation of liquid between two bubble columns connected in an loop configuration, has been demonstrated experimentally, providing data that can be used in the design of such a device. This liquid circulation scheme is potentially useful in chemical looping applications involving aggressive fluids such as molten metals, metal oxides and salts where mechanical pumping is difficult. The observed circulation in the two-column -loop was found to be driven mostly by the difference in bulk density between the fluid in one column and that in the tube connecting it to the adjacent column, while the energy input to each column is attributable mostly to the isothermal expansion of the compressed gas as it rises through the bubble columns. That is, the contribution of kinetic energy from the nozzles is small for the most of the conditions assessed here. The key operational parameters such as gas holdup, the liquid circulation rate and the residence time of the liquid in various zones of the reactor, were found to depend only weakly on the diameter of the single-hole air sparger nozzle for the conditions assessed here, suggesting that turbulence-induced shear in the fluid controls gas holdup. Gas holdup and the other main operational characteristics were also found to be influenced strongly by the superficial air velocity in the riser columns of the reactor. The magnitude of the liquid circulation depends on the total power imparted to the liquid by the expanding gas and increases with the total power input rate. The liquid circulation rate increases with an increase in the height of the gas-free liquid in the columns, in keeping with established theory, but this does not impact significantly the residence time of the liquid in the columns.
21
Acknowledgement This research was performed as part of the Australian Solar Thermal Research Initiative (ASTRI), a project supported by the Australian Government, through the Australian Renewable Energy Agency (ARENA), ARENA 1-SRI002. This program was also supported by the Australian Research Council (ARC) through grant DP180102045. The views expressed herein are not necessarily the views of the Australian Government, and the Australian Government does not accept responsibility for any information or advice contained here. The valuable feedback on the original draft by the anonymous reviewers is gratefully acknowledged.
References Besagni, G., Guédon, G.R., Inzoli, F., 2015. Annular gap bubble column: experimental investigation and computational fluid dynamics modelling. J. Fluids Eng. 138, 011302. Besagni, G., Inzoli, F., 2016. Comprehensive experimental investigation of counter-current bubble column hydrodynamics: Holdup, flow regime transition, bubble size distributions and local flow properties. Chem. Eng. Sci. 146, 259–290. Besagni, G., Inzoli, F., 2017. The effect of liquid phase properties on bubble column fluid dynamics: Gas holdup, flow regime transition, bubble size distributions and shapes, interfacial areas and foaming phenomena. Chem. Eng. Sci. 170, 270–296. Camacho Rubio, F., Sánchez Mirón, A., Cerón García, M.C., García Camacho, F., Molina Grima, E., Chisti, Y., 2004. Mixing in bubble columns: a new approach for characterizing dispersion coefficients. Chem. Eng. Sci. 59, 4369–4376. Chisti, M.Y., 1989. Airlift Bioreactors. Elsevier, London. Chisti, Y., 1998. Pneumatically agitated bioreactors in industrial and environmental bioprocessing: Hydrodynamics, hydraulics and transport phenomena. Appl. Mech. Rev. 51, 33–112.
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Chisti, M.Y., Moo-Young, M., 1987. Airlift reactors: Characteristics, applications and design considerations. Chem. Eng. Commun. 60, 195–242. Chisti, M.Y., Moo-Young, M., 1988. Gas holdup in pneumatic reactors. Chem. Eng. J. 38, 149–152. Chisti, Y., Moo-Young, M., 1993. Improve the performance of airlift reactors. Chem. Eng. Prog. 89(6), 38–45. Chisti, M.Y., Halard, B., Moo-Young, M., 1988. Liquid circulation in airlift reactors. Chem. Eng. Sci. 43, 451–457. Deckwer, W.-D., 1992. Bubble Column Reactors. Wiley, Chichester. Hikita, H., Kikukawa, H., 1973. Gas hold-up in bubble columns: effect of liquid properties. Bull. Univ. Osaka Prefecture Ser. A., Engng. Nat. Sci., 22(2), 151–160. Hikita, H., Asai, S., Tanigawa, K., Segawa, K., Kitao, M., 1980. Gas hold-up in bubble columns. Chem. Eng. J. 20, 59–67. Jafarian, M., Arjomandi, M., Abdollahi, M.R., Nathan, G.J., 2017a. Concentrated solar receiver and reactor systems comprising heat transfer fluid. International Patent Application No. PCT/AU2018/050034
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2017900167 and 2017900564. International Publication No. WO 2018/132875 Al. Jafarian, M., Arjomandi, M., Nathan, G.J., 2017b. Thermodynamic potential of high temperature chemical looping combustion with molten iron oxide as the oxygen carrier. Chem. Eng. Res. Des. 120, 69–81. Jafarian, M., Abdollahi, M.R., Nathan, G.J., 2019. Preliminary evaluation of a novel solar bubble receiver for heating a gas. Solar Energy 182, 264–277. Lee, Y.-K., Ding, S.-Y., Low, C.-S., Chang, Y.-C., Forday, W.L., Chew, P.-C., 1995. Design and performance of an -type tubular photobioreactor for mass cultivation of microalgae. J. Appl. Phycol. 7, 47–51. 23
Merchuk, J.C., 2003. Airlift bioreactors: Review of recent advances. Can. J. Chem. Eng. 81, 324–337. Merchuk, J.C., Siegel, M.H., 1988. Air-lift reactors in chemical and biological technology. J. Chem. Technol. Biotechnol. 41, 105–120. Moo-Young, M., Blanch, H.W., 1981. Design of biochemical reactors: mass transfer criteria for simple and complex systems. Adv. Biochem. Eng. 19, 1–69. Sarafraz, M.M., Jafarian, M., Arjomandi, M., Nathan, G.J., 2017a. Potential use of liquid metal oxides for chemical looping gasification: A thermodynamic assessment. Appl. Energy 195, 702–712. Sarafraz, M.M., Jafarian, M., Arjomandi, M., Nathan, G.J., 2017b. The relative performance of alternative oxygen carriers for liquid chemical looping combustion and gasification. Int. J. Hydrogen Energy 42, 16396–16407. Sarafraz, M.M., Jafarian, M., Arjomandi, M., Nathan, G.J., 2018. Potential of molten lead oxide for liquid chemical looping gasification (LCLG): A thermochemical analysis. Int. J. Hydrogen Energy 43, 4195–4210. Sarafraz, M.M., Jafarian, M., Arjomandi, M., Nathan, G.J., 2019a. Experimental investigation of the reduction of liquid bismuth oxide with graphite. Fuel Process. Technol. 188, 110– 117. Sarafraz, M.M., Jafarian, M., Arjomandi, M., Nathan, G.J., 2019b. The thermo-chemical potential liquid chemical looping gasification with bismuth oxide. Int. J. Hydrogen Energy 44, 8038–8050. Sasaki, S., Uchida, K., Hayashi, K., Tomiyama, A., 2017. Effects of column diameter and liquid height on gas holdup in air-water bubble columns. Exp. Therm. Fluid Sci. 82, 359–366.
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Upham, D.C., Snodgrass, Z.R., Zavareh, M.T., McConnaughy, T.B., Gordon, M.J., Metiu, H., McFarland, E.W., 2017. Molten salt chemical looping for reactive separation of HBr in a halogen-based natural gas conversion process. Chem. Eng. Sci. 160, 245–253. Zhang, T., Wang, T., Wang, J., 2005. Mathematical modeling of the residence time distribution in loop reactors. Chem. Eng. Proc. 44, 1221–1227. Zhang, T., We, C., Ren, Y., Feng, C., Wu, H., 2017. Advances in airlift reactors: modified design and optimization of operation conditions. Rev. Chem. Eng. 33, 163–182.
25
Table 1 Summary of the experimental conditions used.
Gas-free liquid height
Sparger nozzle diameter
Superficial gas velocity (Us, m s1)
(HL, mm)b
(d, mm)b
Column 1
Column 2
1c
300
1.4, 2.0, 2.8, 4.0 and 5.6
0.01 to 0.16
0.01 to 0.16
2
300
2.8
0.039
0 to 0.16
3c
212, 300, 424 and 600
2.8
0.01 to 0.16
0.01 to 0.16
Experimental seta
a In
all cases, the diameters of the columns 1 and 2 were identical at 85 mm.
b These c
variables had identical values in the two bubble columns.
For experimental sets 1 and 3, the superficial gas velocity in the two columns was kept equal for all combinations of sparger nozzle diameters
and heights of the gas-free liquid.
26
Figure captions Fig. 1. Gas-lift reactor with dual interconnected isothermal bubble columns (Jafarian et al., 2017a; Jafarian et al., 2017b). The two bubble columns functioned as riser zones of the circulation loop. Both the risers were injected with a gas. The gas-liquid dispersion leaving the upper zones of the bubble columns was further degassed in gas-liquid separators (degassers) and the remaining liquid was returned to the bottom of the adjacent column. The gas sparged in the bubble columns provided the gas-lift action that moved the liquid between the columns.
Fig. 2. Gas holdup (G) versus superficial gas velocity (Us) in column 1 (a) and column 2 (b). The data are for the experimental set 1 (Table 1). The best fit empirical correlations for gas holdup in the two columns are shown. Identical nozzles of hole diameter d were used in both columns.
Fig. 3. Dependence of the circulation flow rate (Qc) between the bubble columns and the mean residence time (r) of water in each bubble column, on superficial velocity (Us) of gas in the columns. (Us values were identical in the two columns at any time.) Identical nozzles of diameter d were used for gas injection in the two columns. The conditions shown are for experimental set 1 (Table 1). The equation for Qc is based on measured data. The residence time (r) was calculated by substituting the above shown empirical correlation for Qc in Eq. (15).
Fig. 4. Comparison of Eq. (17) (black line) with published correlations (Chisti and MooYoung, 1988; Chisti, 1989; Chisti, 1998) for bubble-flow (red line) and churn-turbulent flow (blue line) regimes for air–water in conventionally operated bubble columns of diverse sizes.
27
Fig. 5. Energy dissipation rate (EF) due to circulation of water between the bubble columns versus superficial gas velocity (Us). The gas in both columns was injected through identical nozzles with a hole diameter d and the Us values were identical. The conditions shown are for experimental set 1 (Table 1). The best fit empirical correlation for the entire data set is shown. EF values were calculated using Eq. (10).
Fig. 6. Dependence of the fractional power input attributable to gas expansion () within the columns, on the superficial gas velocity (Us) in the columns. The gas in both columns was injected through nozzles having a hole diameter d and the Us values were identical in the two columns. The conditions shown are for experimental set 1 (Table 1). values were calculated using Eq. (11).
Fig. 7. Comparison of measured induced liquid circulation rate (Qc) for the modes of operation with: (a) both columns having equal values superficial gas velocities (Us,C1 = Us,C2) at any instance; and (b) column 1 having a fixed aeration rate Us,C1 = 0.039 m s1 for different values of the aeration velocity Us,C2 in column 2. The sparger hole diameter d in both columns was always 2.8 mm and the unaerated liquid height HL was always 300 mm.
Fig. 8. Dependence of the circulation flow rate (Qc) between the bubble columns and the mean residence time (r) of water in each bubble column, on superficial velocity (Us,C2) of the gas injected in column 2. Superficial air velocity (Us,C1) in column 1 remained fixed at 0.039 m s1. Data are shown are for conditions of experimental set 2 (Table 1). Best fit empirical correlations are shown.
28
Fig. 9. Gas holdup (G,C1, G,C2; subscripts C1 and C2 denote columns 1 and 2, respectively) values in the two columns versus superficial gas velocity (Us,C2) in column 2. The superficial gas velocity in column 1 was kept constant at 0.039 m s1. The data are for the conditions in experimental set 2 (Table 1). The best fit empirical equation for gas holdup in column 2 is shown.
Fig. 10. Comparison of the energy dissipation rate (EF) due to circulation of water between the columns, for modes of operation with: (a) column 1 having a fixed aeration rate Us,C1 = 0.039 m s1 for different values of the aeration velocity Us,C2 in column 2; and (b) both columns having equal values superficial gas velocities (Us,C1 = Us,C2) at any instance. The best fit empirical equation for the operation mode (a) is shown. EF values were calculated using Eq. (10). The sparger hole diameter d in both columns was always 2.8 mm and the unaerated liquid height HL was always 300 mm.
Fig. 11. Dependence of the circulation flow rate (Qc) between the bubble columns on superficial gas velocity (Us, identical in both columns) in the columns. Data are shown for identical specified heights (HL) of gas-free water in the two columns (experimental set 3; Table 1). In all case the sparger hole diameter was the same (d = 2.8 mm). Best fit empirical correlations for the data are shown.
Fig. 12. Comparison of the experimentally measured Qc values (Fig. 11) with predictions of Eq. (23). Most of the measured data were within 15% of the solid line, i.e. the line of exact agreement (y = x).
29
Fig. 13. Dependence of the mean residence time (r) of water in each bubble column, on superficial gas velocity (Us) and gas-free liquid height (HL) in the two columns. Both columns had the same values of Us and HL. Data shown are for conditions of experimental set 3 (Table 1).
Fig. 14. Dependence of the gas holdup on superficial gas velocity in bubble column 1 (a) and bubble column 2 (b). Data are shown for various heights (HL; identical in both columns) of gas-free water in the two columns. Superficial gas velocity (Us) was the same in both columns. Data shown are for conditions of experimental set 3 (Table 1). Best fit empirical equations describing the data are shown.
Fig. 15. Dependence of the fractional power input attributable to gas expansion () within the columns, on the superficial gas velocity (Us; identical in both columns) in the columns. Data are shown for various heights (HL; identical in both columns) of gas-free water in the two columns. Data shown are for the conditions of experimental set 3 (Table 1). -values were calculated using Eq. (11). All columns had the same diameter (d = 2.8 mm) of aeration nozzle.
30
Fig. 1
31
Fig. 2
32
Fig. 3
33
G = 49Us0.46 G = 247Us0.97
G (%)
10
G = 63.33Us0.80
Present work (Equation 17) Conventional bubble columns (bubble-flow regime) Conventional bubble columns (churn-turblent flow)
1 0.01
0.1 1
Us (m s )
Fig. 4
34
Fig. 5
35
Fig. 6
36
Fig. 7
37
2.5
110
Qc = 1.02 + 3.06Us,C20.564
100 90
1.5
3
80
r = 105.13 113.39Us,C20.385
1.0
70 60
0.5
Qc
50
r 0.0 0.00
0.02
0.04
0.06
0.08
0.10 1
Us,C2 (m s )
Fig. 8
38
0.12
0.14
0.16
40 0.18
r (s)
1
Qc (dm min )
2.0
14
G,C1, G,C2 (%)
12
Column 1 (Us,C1 = 0.039 m s1) Column 2
10
G,C2 = 67.57Us,C20.841
8 6 4
G,C1
2 0 0.00
0.02
0.04
0.06
0.08
0.10 1
Us,C2 (m s )
Fig. 9
39
0.12
0.14
0.16
0.18
0.05
EF (W)
0.04
(b) Us,C1 = Us,C2
0.03
(a) 0.02
0.01
EF = 2.3x103 + 0.145Us,C2 + 0.121Us,C22 0.00 0.00
0.02
0.04
0.06
0.08
0.10
0.12
Us,C2 (m s1)
Fig. 10
40
0.14
0.16
0.18
0.20
10
Qc = 10.43Us
0.457
3
0.534
Qc = 7.03Us
1
Qc (dm min )
Qc = 9.84Us
1
0.501
Qc = 4.70Us0.469 HL (mm) 600 424 300 212
0.1 0.01
0.1
1 1
Us (m s )
Fig. 11
41
5
4
15%
3
1
Predicted Qc (dm min )
+15%
3
2
HL (mm)
212 300 424 600
1
0
0
1
2
3
4 3
1
Experimental Qc (dm min )
Fig. 12
42
5
Fig. 13
43
Fig. 14
44
Fig. 15
45
Highlights A novel two-column gas-lift reactor configuration was developed. Reactor comprised two interconnected riser zones forming an -shaped loop. Mixing and circulation were achieved solely by a gas-lift action. Possible applications in chemical looping with different gases in the two columns. Correlations were developed for gas holdup and liquid circulation rate.
46
CRediT authorship contribution statement Mehdi Jafarian: Conceptualization, Investigation, Methodology, Data Curation, Formal analysis, Validation, Visualization, Writing - Original Draft. Yusuf Chisti: Investigation, Visualization, Formal analysis, Writing-review& editing. Graham J. Nathan: Conceptualization, Supervision, Formal analysis, Methodology, Writing review & editing, Supervision, Funding acquisition.
47
S0009-2509(20)30106-8 https://doi.org/10.1016/j.ces.2020.115574 CES 115574
To appear in:
Chemical Engineering Science
Received Date: Revised Date: Accepted Date:
25 October 2019 7 February 2020 11 February 2020
Please cite this article as: M. Jafarian, Y. Chisti, G.J. Nathan, Gas-lift circulation of a liquid between two interconnected bubble columns, Chemical Engineering Science (2020), doi: https://doi.org/10.1016/j.ces.2020.115574
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© 2020 Published by Elsevier Ltd.
Gas-lift circulation of a liquid between two inter-connected bubble columns
Mehdi Jafarian a,*, Yusuf Chisti b, Graham J. Nathan a
a
Centre for Energy Technology, School of Mechanical Engineering, The University of
Adelaide, SA 5005, Australia
b
School of Engineering, Massey University, Private Bag 11 222, Palmerston North, New
Zealand
* Corresponding author: School of Mechanical Engineering, The University of Adelaide, Adelaide, SA 5005, Australia. Tel.: +61-8-8303-5460; E-mail: [email protected] (M. Jafarian).
1
Abstract A pneumatically-mixed novel configuration of two inter-connected bubble column reactors is reported using air-water to demonstrate the device, enabling circulation of a liquid between the columns solely by a gas-lift mechanism. The new reactor may be used for circulating a high temperature/corrosive liquid between the columns, enabling the cyclical processing of a liquid medium with two different gaseous reactants. The effects of superficial gas velocity, unaerated liquid height in the columns and the diameter of aeration nozzles are discussed with reference to the gas holdup in, and rate of liquid circulation between, the two columns, together with the energetics of circulation. The height of the gas-free liquid in the reactor was found to influence the liquid circulation velocity positively, but the effect of aeration nozzles diameter was marginal. The lowest and highest measured flow rates of circulating water between the columns were 0.7 dm3 min1 and 4.2 dm3 min1, respectively.
Keywords: airlift reactor; bubble column; chemical looping; two-phase flow; gas holdup.
2
Nomenclature A
cross sectional area of a column (m2)
AC1
cross sectional area of column 1 (m2)
AC2
cross sectional area of column 2 (m2)
a
parameter in Eq. (20) (dm3 min1)
aL
gas–liquid interfacial area per unit volume of liquid (m1)
b
parameter in Eq. (20) (-)
d
internal diameter (ID) of gas injection nozzle (m)
EF
frictional energy dissipation rate in the entire circulation loop (W)
EK,C1 kinetic energy input rate in column 1 (W) EK,C2 kinetic energy input rate in column 2 (W) ES
energy dissipation rate due to stagnant gas in downflow ducts (W)
EW,C1 energy dissipation rate in the liquid wakes behind the bubbles in column 1 (W) EW,C2 energy dissipation rate in the liquid wakes behind the bubbles in column 2 (W) EX,C1
energy input rate due to isothermal gas expansion in column 1 (W)
EX,C2
energy input rate due to isothermal gas expansion in column 2 (W)
g
gravitational acceleration (m s2)
HL
height of gas-free liquid in a bubble column (m)
Hm1
manometer height shown in Fig. 1 (m)
Hm2
manometer height shown in Fig. 1 (m)
ΔHm
manometer reading (m)
hD
height of gas-liquid dispersion in a column (m)
hD,C1
height of gas-liquid dispersion in column 1 (m)
hD,C2
height of gas-liquid dispersion in column 2 (m)
mG
mass flow rate of gas in a column (kg s1) 3
Ph
headspace pressure (Pa)
Qc
flow rate of the circulating liquid between bubble columns (dm3 min1)
Qm
molar flow rate of air in the column (kmol s1)
R
gas constant (kJ kmol1 K1)
T
absolute temperature (K)
UL
superficial liquid velocity in bubble column (m s1)
UL,C1 superficial liquid velocity in column 1 (m s1) UL,C2 superficial liquid velocity in column 2 (m s1) Us
superficial air velocity in bubble column (m s1)
Us,C1
superficial air velocity in column 1 (m s1)
Us,C2
superficial air velocity in column 2 (m s1)
un
gas velocity at the injection nozzle (m s1)
VG
volume of the gas phase (m3)
VL
volume of the liquid phase (m3)
VL,C1
volume of liquid in column 1 (m3)
VL,C2
volume of liquid in column 2 (m3)
Vt
total volume of gas-liquid dispersion in an aerated column (m3)
Greek letters
G
gas holdup (-)
G,C1
gas holdup in column 1 (-)
G,C2
gas holdup in column 2 (-)
contribution of gas expansion to total power input (%)
ρL
density of the liquid (kg m3)
4
ρG
density of gas (kg m3)
fraction of jet’s momentum energy transferred to liquid (-)
r
mean residence time of the liquid in a bubble column (s)
5
1. Introduction Gas-lift reactors are a well-established technology in gas-liquid contacting operations, employing the action of a compressed gas to achieve good mixing and high rates of heat and mass transfer in a liquid or slurry (Merchuk and Siegel, 1988; Chisti, 1998; Merchuk, 2003; Zhang et al., 2017). The avoidance of moving parts, such as shafts and stirrers, reduces capital and operational costs (Chisti and Moo-Young, 1987; Jafarian et al., 2019) and, in biological processes, reduces the potential for contaminating microorganisms entering the vessel through shaft seals. Conventional gas-lift reactors, such as those used in biological treatment of wastewater and other bioprocessing applications, typically contact only a single gas (e.g. air) with the liquid, or slurry, being mixed. However, much less attention has been paid to bubbling reactors that enable the looping of reactants between two vessels. Various geometric configurations of multi-column gas-lift reactors have been described previously (Chisti and Moo-Young, 1987; Chisti, 1998; Merchuk, 2003). In all cases, the reactor consists of two relatively tall vertical zones that are linked near to the top and bottom. One of these vertical zones (the riser) is typically bubbled with a gas. This results in an up flow of the gas-liquid dispersion in the riser zone and a downflow in the interconnected vertical downcomer zone that is mostly free of gas (Merchuk and Siegel, 1988; Merchuk, 2003). This circulatory flow is driven mostly by the difference in bulk densities between the gas-liquid dispersion in the riser and that in the downcomer zones (Chisti and Moo-Young, 1987). In specific cases, the momentum of the compressed gas injected near to the bottom of the riser zone may also contribute to circulation of the fluid. In most of these gas-lift reactors, the configuration of the top zone where the riser and the downcomer connect, ensures that most of the gas disengages from the fluid before it enters the downcomer. Gas disengagement is necessary for achieving high rates of liquid circulation. Often the top zone is purpose-designed to enhance gas-liquid separation (Chisti and Moo-Young, 1993; Chisti, 1998). Thus, while this
6
system is suitable for performing a single reaction between a gas and liquid or slurry, by itself it cannot be used to perform two cyclical reactions as is needed for chemical looping processes. To address the above gap, a novel gas-lift configuration capable of achieving this has been recently developed by Jafarian et al. (2017a; 2017b). This device uses dual interconnected gas-lift reactors (Fig. 1) to circulate a liquid between two reactors, each bubbling a different gaseous reactants, to be suitable for chemical looping processes. As shown in Fig. 1, this novel configuration is comprised of two interconnected gas-sparged bubble column riser zones. The fluid leaving the top of each column passes through degassing zones to enter the other column near the bottom to produce an -shaped flow loop (Fig. 1). The degassing zones also prevent mixing between the different gaseous reactants by assuring disengagement of the bubbles from the liquid exiting a bubble column and before entry to the other column. Two different reactant gases can therefore be used simultaneously without being mixed. The absence of moving parts in the reactor (Fig. 1) is expected to improve its utility under challenging processing conditions involving high temperatures and corrosive reactants that can be difficult to circulate with mechanical pumping. Examples include molten metal oxides at high temperature, which are used as oxygen carriers in liquid chemical looping in combustion/gasification systems (Jafarian et al., 2017b; Sarafraz et al., 2017a; Sarafraz et al., 2017b; Sarafraz et al., 2018). Similarly, the circulation of molten bromide salt is used in a chemical looping process for oxidation of hydrogen bromide to bromine (Upham et al., 2017). Although airlift-driven tubular-loop photobioreactors conceptually similar to that shown in Fig. 1, have been reported previously (Lee et al., 1995), no data are presently available on their operation or hydrodynamic characteristics (Fig. 1). Therefore, the first objective of the present work was to meet this need. Although the factors that influence circulation rate of liquids in airlift reactors are well known for single riser configurations, their relative significance in gas-lift reactors with dual 7
inter-connected riser zones (Fig. 1), is not. Liquid circulation rate affects the residence time of the fluid in different zones of a reactor such as the one shown in Fig. 1 and, hence, the kinetics of any reactions being carried out (Zhang et al., 2005). For example, the degree of oxidation of fuel in chemical looping combustion and gasification systems has been shown to depend on the recirculation flow rate of the molten oxygen carrier between the fuel and air compartments of the reactor (Jafarian et al., 2017b; Sarafraz et al., 2017a; Sarafraz et al., 2017b; Sarafraz et al., 2018; Sarafraz et al., 2019a; Sarafraz et al., 2019b). On this basis, the aim of the present investigation is to advance the understanding of the factors that control the circulation rate in a dual interconnected bubble column system and to provide an experimental assessment of the hydrodynamics and liquid circulation characteristics of the reactor configuration shown in Fig. 1.
2. Methodology 2.1. Theoretical analysis All energy for circulation of the liquid in the two-column -loop reactor system (Fig. 1) is generated by the gas injected into the two bubble columns. Part of the energy input is a consequence of the work done during isothermal expansion of the gas as it moves up the columns, while the remainder is a consequence of the kinetic energy of the gas jets issuing from the sparger nozzles. Energy is dissipated in the wakes of the circulating liquid behind the bubbles (Chisti et al., 1988; Chisti, 1989) that rise in the bubble columns and from the changes in cross section of the flow path and friction in the flow circuit including the connecting pipes and the gas disengaging devices (Fig. 1). An energy balance over the circulation loop (Fig. 1) can be written as follows: 𝐸𝑋,𝐶1 + 𝐸𝑋,𝐶2 + 𝐸𝐾,𝐶1 + 𝐸𝐾,𝐶2 = 𝐸𝑊,𝐶1 + 𝐸𝑊,𝐶2 + 𝐸𝐹 + 𝐸𝑆
8
(1)
In the above equation: EX,C1 and EX,C2 are the energy input rates due to isothermal expansion of the gas in columns 1 and 2, respectively; EK,C1 and EK,C2 are the kinetic energy input rates due to gas injected at the bottoms of columns 1 and 2, respectively; EW,C1 and EW,C2 are energy dissipation rates in the liquid wakes behind the bubbles rising in columns 1 and 2, respectively; EF is the energy dissipation rate due to cross sectional changes and friction in the circulation loop; and ES is the energy dissipation rate due to any stagnant gas (Chisti et al., 1988; Chisti, 1989) in the two downflow ducts connecting the bubble columns (Fig. 1). Equation (1) disregards the kinetic energy dissipation rate at the top openings of the bubble columns because of the low gas flow velocity in those regions. It also ignores frictional losses in the bubble columns because of the low liquid flow rates in them. The ES term in Eq. (1) can be neglected as no stagnant gas bubbles were observed in the interconnecting ducts because of the disengagement of gas in the gas-liquid separators (Fig. 1). The energy dissipation rate in the wakes behind the bubbles in column 1 (EW,C1) and column 2 (EW,C2) of the circulation loop was calculated using the following equations (Chisti et al., 1988): 𝐸𝑊,𝐶1 = 𝐸𝑋,𝐶1 + 𝐸𝐾,𝐶1 ― 𝜌𝐿𝑔𝑈𝐿,𝐶1𝜀𝐺,𝐶1ℎ𝐷,𝐶1𝐴𝐶1
(2)
𝐸𝑊,𝐶2 = 𝐸𝑋,𝐶2 + 𝐸𝐾,𝐶2 ― 𝜌𝐿𝑔𝑈𝐿,𝐶2𝜀𝐺,𝐶2ℎ𝐷,𝐶2𝐴𝐶2
(3)
where L is the density of the liquid, g is gravitational acceleration, UL is the superficial liquid velocity, G is the gas holdup, hD is the height of gas-liquid dispersion in the column, and A is the cross sectional area of the column. The subscripts C1 and C2 denote columns 1 and 2, respectively. The total volume Vt of gas-liquid dispersion in an aerated column can be related to the height of dispersion (hD) in the column and the cross sectional area (A), as follows: (4)
𝑉𝑡 = ℎ𝐷𝐴 where Vt is the sum of the volumes of the gas (VG) and liquid phases (VL), or: 9
(5)
𝑉𝑡 = 𝑉𝐺 + 𝑉𝐿
Furthermore, by definition, the gas holdup (G), or volume fraction of gas in gas-liquid dispersion, depends on the volumes of the individual phases, as follows: 𝜀𝐺 =
𝑉𝐺
(6)
𝑉𝐺 + 𝑉𝐿
From Eq. (5) and Eq. (6), the following equation can be obtained:
(
𝑉𝑡 = 𝑉𝐿
)
1 1 ― 𝜀𝐺
(7)
Substitution of Eq. (4) and Eq. (7) into Eq. (2), leads to the following: 𝐸𝑊,𝐶1 = 𝐸𝑋,𝐶1 + 𝐸𝐾,𝐶1 ― 𝜌𝐿𝑔𝑈𝐿,𝐶1𝑉𝐿,𝐶1
(
𝜀𝐺,𝐶1
)
(8)
)
(9)
1 ― 𝜀𝐺,𝐶1
Analogously, for column 2 we obtained: 𝐸𝑊,𝐶2 = 𝐸𝑋,𝐶2 + 𝐸𝐾,𝐶2 ― 𝜌𝐿𝑔𝑈𝐿,𝐶2𝑉𝐿,𝐶2
(
𝜀𝐺,𝐶2
1 ― 𝜀𝐺,𝐶2
With ES = 0, as discussed above, the substitution of Eq. (8) and Eq. (9) into Eq. (1), provides the following expression for energy dissipation rate (EF) in the loop:
[
𝐸𝐹 = 𝜌𝐿𝑔 𝑈𝐿,𝐶1𝑉𝐿,𝐶1
(
𝜀𝐺,𝐶1
)+𝑈
1 ― 𝜀𝐺,𝐶1
)]
(
𝜀𝐺,𝐶2 𝐿,𝐶2𝑉𝐿,𝐶2 1 ― 𝜀𝐺,𝐶2
(10)
The percentage () of the total input power that is transferred to gas expansion were calculated as follows: 𝐸𝑋,𝐶1 + 𝐸𝑋,𝐶2
(11)
𝜃 = 𝐸𝑋,𝐶1 + 𝐸𝑋,𝐶2 + 𝐸𝐾,𝐶1 + 𝐸𝐾,𝐶2 × 100
In the above equation, the power input resulting from gas expansion (i.e. EX,C1, EX,C2) and the kinetic power input (i.e. EK,C1, EK,C2), were calculated using the following equations (Chisti, 1989): 𝐸𝑋 = 𝜌𝐿𝑔𝑈𝑠𝑉𝐿
(12)
𝐸𝐾 = 0.5 𝛺 𝑚𝐺𝑢2𝑛
(13)
where 𝑚𝐺 is the mass flow rate of gas in a column, un is the gas velocity at the injection nozzle and is the fraction of the jet’s kinetic energy that is transferred to the liquid. Based on earlier 10
work, is estimated to be approximately 0.06 (Moo-Young and Blanch, 1981; Chisti, 1989). The superficial gas velocity in the above equations was calculated using the following equation (Chisti, 1989): 𝑄𝑚𝑅𝑇
(
𝑈𝑠 = 𝑉𝐿𝜌𝐿𝑔𝑙𝑛 1 +
)
𝜌𝐿𝑔𝐻𝐿 𝑃ℎ
(14)
where Qm is the molar flow rate of air, R is the gas constant, T is the absolute temperature, VL is the volume of liquid in a column, and Ph is the headspace pressure in the columns open to atmosphere. The mean residence time (r) of the liquid in each bubble column, was calculated using the following equation (Merchuk, 2003; Zhang et al., 2017): 𝜏𝑟 =
𝑉𝐿
(15)
𝑄𝑐
where Qc is the flow rate of the circulating liquid between the bubble columns. The reactor shown in Fig. 1 can potentially be operated in two alternative modes: either (1) with equal values of superficial air velocities in the two columns, or (2) with different values of superficial gas velocity in the two columns. Mode 2 may be useful if different gaseous reactants are to be used, for example (Jafarian et al., 2017b; Sarafraz et al., 2017a; Sarafraz et al., 2017b; Sarafraz et al., 2018). Both of these operating modes were investigated, as is discussed in the following sections.
2.2. Experimental reactor and measurements The reactor comprised of two identical cylindrical bubble columns (diameter = 85 mm) connected to each other as shown in Fig. 1. Liquid circulation and mixing in the reactor were studied using a deionized water–air system. Air from a compressor (Fig. 1) was sparged at the bottom of the bubble columns using identical nozzles of specified diameter. Two electronic
11
mass flow controllers were used to control the flow rate of air to the bubble columns at specified values. The air flow rate was reproducible to within 2% of the set value. Two ultrasonic flow meters (accurate to 5% of the measured value in the range of 0.2– 0.5 dm3 min1 and 3% of the measured value in the range of 0.5–25 dm3 min1) were used in parallel to measure the flow rate of the liquid flowing between the columns (Fig. 1; only one flow meter is shown for clarity). The steady-state flow was measured at 0.1 s intervals for 60 s and the data were averaged. LabVIEW software (National Instruments, Austin, TX, USA) was used to post-process the data. Gas holdup (G), or volume fraction of gas in a gas-liquid dispersion, is an important operational characteristic of gas-lift reactors for several reasons: (1) a gas-liquid reactor must be sized to accommodate the expected gas holdup (Chisti, 1989); (2) differences in gas holdup in different zones of a gas-lift reactor determine the rate of liquid circulation in the reactor loop (Chisti et al., 1988; Chisti, 1989); (3) absorption, or mass transfer, of a gaseous reactant is strongly influenced by the gas–liquid interfacial area (aL) per unit volume of liquid in the reactor (Chisti, 1989) so that aL depends on both the gas holdup and the average bubble diameter; and (4) the average residence time of gas in liquid depends on gas holdup (MooYoung and Blanch, 1988; Chisti, 1989). The gas holdup (G) was measured in both bubble columns using the manometeric method (Chisti, 1989). Two manometers were connected to the columns as shown in Fig. 1. The manometer readings (ΔHm) were recorded digitally with a Sony ILCE-6000 camera. These images were then post processed to obtain measurements of ΔHm. The ΔHm value was reproducible to within 1% of the measured value near the upper limit and to within 12% at the low end of the measured values. The average steady-state gas holdup in the column between the taps of the manometer was calculated using the following equation (Chisti, 1989):
12
𝜀𝐺 =
(
𝜌𝐿
)
𝜌𝐿 + 𝜌𝐺
𝐻𝑚2 ― 𝐻𝑚1 𝐻𝑚
=
(
𝜌𝐿
)
𝛥𝐻𝑚
(16)
𝜌 𝐿 + 𝜌 𝐺 𝐻𝑚
where ρL was the density of the liquid, ρG was the density of gas, ΔHm was the manometer reading (Hm = Hm2 – Hm1; Fig. 1) and Hm was the distance between the taps of the manometer. The experimental variables were the superficial gas velocity (Us) in the bubble columns, the height (HL) of the gas-free liquid in each column and the diameter (d) of the gas injection nozzles. The unaerated liquid height HL was identical in the two bubble columns and was level with the height of the horizontal exit port located near to the top of each column (Fig. 1). In the various experiments, HL ranged from 0.212 m to 0.600 m. Whenever HL was changed, the location of the fluid exit port in the two columns was also changed. Identical gas injection nozzles were used in the two bubble columns for each experiment, although the internal diameter of the nozzles was varied (d = 1.4, 2.0, 2.8, 4.0 and 5.6 mm). All other geometric measurements of the reactor system remained fixed as shown in Fig. 1. The specific combinations of HL, sparger nozzles and aeration velocities used in different experiments are summarized in Table 1. Experiments were carried out at room temperature (24 C) and atmospheric pressure, so that the correlations developed in this work apply to operation at atmospheric pressure.
3. Results and discussion 3.1. Operation with equal aeration velocity in the two bubble column zones Figure 2 shows that the gas holdup in the two bubble column zones depends on the superficial air velocity via a power law. This is consistent with other well-known observations about gas holdup in conventional bubble columns (Hikita et al., 1980; Chisti and Moo-Young, 1988; Chisti, 1989; Deckwer, 1992; Besagni et al., 2015; Besagni and Inzoli, 2016; Besagni and Inzoli, 2017). Similar values of g were measured in both columns (Fig. 2), which is expected because the columns were identical and operated under the same conditions. The gas 13
holdup correlations for the two columns are also identical to within measurement accuracy (Fig. 2), giving confidence in the measurements. Therefore, the two correlations were combined into the following equation: 𝜀𝐺 = 63.33 𝑈0.80 𝑠
(17)
Equation (17) was obtained by regression of the combined data for the two columns. Gas holdup showed barely any dependence on the diameter of the single-hole sparger nozzle, notwithstanding the broad range of hole diameters (d = 1.4 to 5.6 mm) tested (Fig. 2). This is consistent with observations in conventional bubble columns (Hikita et al., 1973; Hikita et al., 1980), suggesting that the gas holdup is controlled by the prevailing turbulence level in the water rather than by the initial size of the gas bubbles generated at the sparger nozzle (Chisti, 1989; Deckwer, 1992). The bubble columns in the present work (Fig. 1) had a superimposed liquid up-flow because each column also served as the riser zone of the airlift circulation loop (Fig. 1). This is unlike a conventional bubble column, which lacks a net up-flow of liquid (Chisti, 1989; Camacho Rubio et al., 2004). The measured liquid circulation rate between the columns is shown in Fig. 3. The rate of the liquid circulation and up flow of liquid in the bubble columns can be seen to increase with increasing aeration velocity (Fig. 3). An imposed cocurrent liquid flow in a bubble column reduces the gas holdup relative to the case of an otherwise identical conventional column operated without flow through of liquid (Chisti, 1989). In addition, presence of flowing liquid in a bubble column postpones, or suppresses, the transition of flow regime from the bubble-flow regime to churn-turbulent (or heterogeneous) flow regime, to higher values of the superficial gas velocity. In air–water in conventional bubble columns this flow transition occurs around a superficial gas velocity value of 0.05 m s–1 (Chisti, 1989; Chisti, 1998).
14
Figure 4 compares the gas holdup predictions of Eq. (17) with two well-known equations developed for gas holdup in conventional bubble columns operated with air–water (Chisti and Moo-Young, 1988; Chisti, 1989; Chisti, 1998). As explained above, the gas holdup in the present study was lower and the flow transition from the bubble-flow regime to churnturbulent regime did not occur at the superficial aeration velocity expected for a conventional column (Fig. 4). The two literature equations given in Fig. 4 for the bubble-flow and churnturbulent flow regimes in air–water, have correlated data from many different conventionally operated bubble columns. For the columns in the present work (Fig. 3), the liquid circulation rate was 0.7 dm3 min–1 for the lowest air flow velocity (0.01 m s1), while it was 2.8 dm3 min–1 at the highest value of aeration velocity (0.16 m s1). This range of liquid volume flow rates corresponds to a liquid superficial velocity range in a bubble column of 2.1103 to 9.6103 m s1. For fixed geometry of the circulation loop (Fig. 1) and complete disengagement of gas from the interlinking tubes, the gas holdup in the risers (i.e. the bubble columns) provides the main driving force for liquid circulation. Therefore, as the gas holdup increased with increasing aeration rate in the columns, the liquid flow velocity increased, leading to a decreasing residence time of liquid in each column (Fig. 3). Over the operational range of aeration velocities (around 0.01–0.16 m s1), the residence time varied by around 3.7-fold (from 149 to 40 s). In principle, it would be possible to have different liquid residence times in the two columns, if the columns had different diameters but were otherwise identical. As gas holdup was independent of the diameter of the aeration nozzle, barely any effect of the nozzle diameter could be seen on the induced liquid circulation rate and the liquid residence time in a column (Fig. 3). A lack of significant effect of aeration nozzle diameter on liquid circulation velocity also suggests that the kinetic energy of the air jet issuing from the sparger contributes little to liquid circulation for these conditions. 15
The energy dissipation rate (EF), calculated using the measured values of gas holdup in Eq. (10), is shown in Fig. 5 for the range of superficial air velocities in the two columns. It can be seen that EF is independent of diameter of the aeration nozzle as is the gas holdup in Fig. 2. Also EF dependence on the superficial aeration velocity is well described by an empirical power law, as follows: 𝐸𝐹 = 0.395 𝑈1.18 𝑠
(18)
This equation is consistent with Eq. (10) if the dependence of gas holdup on the superficial aeration velocity is considered. The power imparted to the liquid was contributed by isothermal expansion of the air as it rose up the bubble columns and the kinetic energy of the air jets formed at the sparger nozzles. Fig. 6 presents (calculated from Eq. 11) as a function of the superficial gas velocity. This shows that, for all but the two smallest nozzles (d 2.0 mm), 80% or more of the energy input to the liquid in the riser column is provided by isothermal expansion of the air. However, for the narrow diameter nozzles operated at high aeration rates, the contribution of the kinetic energy of the gas jet becomes significant, showing that it is possible to develop devices that operate in this regime. Nevertheless, for all cases investigated here, the contribution of gas expansion to total energy input is always at least 50%. As expected, the kinetic energy input increases as the cross sectional area of the nozzle is reduced for a given superficial gas velocity.
3.2. Operation with different aeration rates in the bubble column zones The measured liquid circulation rate (Qc) between the two bubble columns is compared in Fig. 7 for the two different modes of operation: (a) the mode with identical values of the superficial air velocities in both columns (i.e. Us,C1 = Us,C2) and (b) the mode with the value of Us,C1 fixed at 0.039 m s1 for different Us,C2 values in column 2 (Table 1). The case with Us = 0 m s1 in column 2 is equivalent to operating as a conventional airlift reactor with a single 16
aerated riser zone, despite the unusual -shaped recirculation loop. For this comparison, both column were installed with an identical sparger nozzle (d = 2.8 mm) and the height of gas-free liquid in both columns was also fixed at HL = 300 mm. The two columns have an identical energy input when the superficial aeration velocity in column 2 is 0.039 m s1 (Fig. 7). Fig. 7 reveals that, irrespective of the mode of operation, the induced liquid circulation rate depends on the total energy input to the reactor system. For Us,C2 = 0.039 m s1, the energy input rate due to aeration in both modes of operation is equal, as is the liquid circulation rate (Fig. 7). For Us,C2 > 0.039 m s1, the reactor system with both the bubble columns aerated at equal aeration velocities always has a higher total energy input rate (black line in Fig. 7) and, consequently also, a higher liquid circulation rate than the reactor loop operated with Us,C1 fixed at 0.039 m s1 (red line in Fig. 7). The situation is reversed for Us,C2 < 0.039 m s1, so that the total power input rate in reactor system (i.e. the red line in Fig. 7) is greater than the case with Us,C1 = 0.039 m s1 (i.e. the black line in Fig. 7). For all cases, the total liquid circulation flow rate increases with the total energy input. Fig. 8 presents both the liquid circulation rate (Qc) and the mean residence time of water in each column (r) as a function of the superficial aeration velocity in column 2 (Us,C2). It can be seen that a minimum rate of circulation occurs with no gas injection in column 2 (i.e. Us,C2 = 0 m s1) because of the constant aeration through column 1 (Us,c1 = 0.039 m s1). This case also corresponds to the maximum residence time through the two columns, as expected. For non-zero values of Us, both the total circulation rate and the residence time are well described by empirical correlations, so that Qc = 1.02 + 3.06Us,C20.564 and r = 105.13 113.39Us,C20.385. The values of gas holdup corresponding to the data in Fig. 8, in the two columns are presented in Fig. 9. Gas holdup in column 2 increases with increasing superficial air velocity Us,C2 in keeping with the dependence of this variable on aeration velocity discussed above. Similarly, the gas holdup remains constant at 5% in column 1, as consistent with the constant 17
aeration velocity in that column (Us,C1 = 0.039 m s1). Nonetheless, both columns contribute to the flow circulation, as discussed above. Maintaining different values of gas holdup in the two columns may be desirable for some applications, such as where two gas-liquid reactions are being carried out in these zones for chemical looping applications. The energy loss due to circulation of water for the mode of operation with unequal aeration rates in the two columns can be seen to depend on the aeration velocity (Us,C2) in column 2 as shown in Fig. 10. The EF values, as calculated using Eq. (10), are well described by the following empirical equation (Fig. 10): 𝐸𝐹 = 2.3 × 10 ―3 +0.145𝑈𝑠,𝐶2 +0.121𝑈2𝑠,𝐶2
(19)
The non-zero value of energy loss for Us,C2 = 0 m s1 (Eq. 19) can be attributed to the circulation induced by the constant aeration rate in column 1 (Us,C1 = 0.039 m s1). Relative to operation with equal superficial air velocities in the two columns (operation mode (b), Fig. 10), the mode of operation with unequal aeration has lower values of EF where Us,C2 > 0.039 m s1, but higher values otherwise. This behavior is also explained by the differences in the induced liquid flow rate in different modes of operation discussed above.
3.3. Operation with equal aeration velocity in both bubble columns: effects of the height of gas-free liquid The dependence of the circulation flow rate (Qc) between the bubble columns on superficial gas velocity for various heights of unaerated liquid in the columns is presented in Fig. 11. In any given experiment, the HL values in the two columns were identical and varied over a range of 212 to 600 mm. The measured liquid circulation flow rate (Qc) increased with the aeration velocity (Us) and HL, consistent with previous work (Chisti et al., 1988; Chisti, 1989; Chisti and Moo-Young, 1993).
18
The observations in Fig. 11 were are also consistent with those seen previously in Fig. 3 for the same mode of operation but with a specific value of HL (300 mm). The dependence of the liquid circulation rate on the aeration velocity in the two columns is well described by the following power law: 𝑄𝑐 = 𝑎 𝑈𝑏𝑠
(20)
Here, the dependence of the values of the empirical constants a and b are as follows: 𝑎 = 4.352𝑙𝑛 𝐻𝐿 +11.852
(21)
𝑏 = 0.4 𝐻𝐿―0.142
(22)
Combining these dependences, Eq. (20) can be rewritten as follows: ―0.142
𝐻𝐿 𝑄𝑐 = (4.352𝑙𝑛 𝐻𝐿 + 11.852)𝑈0.4 𝑠
(23)
This empirical equation was found to correlate all of the data in Fig. 11 except for 2 data points to within 15% of an exact agreement, as is shown in Fig. 12. The correlation is consistent with earlier studies in conventional airlift reactors, which show that the rate of induced liquid circulation increases with the height of dispersion in the reactor, for otherwise identical values of aeration rate and other geometric characteristics of the reactor (Chisti et al., 1988; Chisti, 1989). Fig. 13 presents the values of r calculated using Eq. (15) as a function of the aeration rate (identical in the two columns) for various values of HL (identical in both columns). A strong inverse relationship can be seen. However, any effect of HL on the liquid residence time is small, unlike the influence of HL on the liquid circulation rate (Fig. 11). This can be explained by the co-dependence on HL of Qc and the volume of liquid (VL in Eq. 15). Although Qc increases with increasing HL, so does VL. So that the net effect on the residence time of changing HL is minor (Fig. 13). The dependence of gas holdup on superficial velocity is presented for the two columns in Fig. 14. The response is almost identical, consistent with nominally identical values of HL 19
and aeration velocities, with the differences attributable to experimental uncertainty. The gas holdup can be seen to increase with increasing aeration velocity for a given value of unaerated liquid height in a column, as explained above. However, at any fixed aeration velocity, increasing the height of the gas-free liquid increases the gas holdup (Fig. 14). This can be possibly attributed to an increase in the liquid superficial velocity with an increase in HL (as seen in Fig. 11) which, in turn, may increase the turbulence-induced shear rate within the liquid. Such an increase in turbulent shear can be expected to reduce the mean bubble size and, therefore, also reduce the average rise velocity of the bubbles. In conventional bubble columns without an imposed liquid flow, neither the height of liquid in the column nor the column diameter influence the average gas holdup (Chisti, 1989; Sasaki et al., 2017), provided that the columns are sufficiently large (e.g. diameter 0.1 m, gasfree liquid height 1 m). In contrast, the height of gas-lift reactors profoundly affects the gas holdup and liquid circulation (Chisti et al., 1988; Chisti, 1989; Chisti and Moo-Young, 1993). In airlift reactors, the efficiency of gas-liquid separation in the top region of the reactor as well as the geometry of the degassing zone (Chisti, 1998; Chisti and Moo-Young, 1993) further impact liquid circulation rate and gas holdup (Chisti et al., 1988; Chisti, 1989; Chisti and MooYoung, 1993). The dependence on superficial gas velocity of the fraction of the total energy input () that was contributed by isothermal expansion of the gas is presented in Fig. 15 for a series of heights of gas-free liquid. While the general trend is similar to that in Fig. 6, it can be seen here that an increase in HL causes an increase in for all values of Us. This can be attributable to the increase in hydrostatic pressure at the bottom of the column with an increase in HL, meaning that the gas needs to expand from a higher pressure at the bottom of the column to reach the atmospheric pressure at the surface of the gas-liquid interface. It is also noteworthy that, with
20
the exception of a single case (the -value at HL = 212 mm and Us 0.16 m s1; Fig. 15), 80% of the total power input to the liquid is attributable to the isothermal expansion of air.
4. Conclusions A gas-lift driven circulation of liquid between two bubble columns connected in an loop configuration, has been demonstrated experimentally, providing data that can be used in the design of such a device. This liquid circulation scheme is potentially useful in chemical looping applications involving aggressive fluids such as molten metals, metal oxides and salts where mechanical pumping is difficult. The observed circulation in the two-column -loop was found to be driven mostly by the difference in bulk density between the fluid in one column and that in the tube connecting it to the adjacent column, while the energy input to each column is attributable mostly to the isothermal expansion of the compressed gas as it rises through the bubble columns. That is, the contribution of kinetic energy from the nozzles is small for the most of the conditions assessed here. The key operational parameters such as gas holdup, the liquid circulation rate and the residence time of the liquid in various zones of the reactor, were found to depend only weakly on the diameter of the single-hole air sparger nozzle for the conditions assessed here, suggesting that turbulence-induced shear in the fluid controls gas holdup. Gas holdup and the other main operational characteristics were also found to be influenced strongly by the superficial air velocity in the riser columns of the reactor. The magnitude of the liquid circulation depends on the total power imparted to the liquid by the expanding gas and increases with the total power input rate. The liquid circulation rate increases with an increase in the height of the gas-free liquid in the columns, in keeping with established theory, but this does not impact significantly the residence time of the liquid in the columns.
21
Acknowledgement This research was performed as part of the Australian Solar Thermal Research Initiative (ASTRI), a project supported by the Australian Government, through the Australian Renewable Energy Agency (ARENA), ARENA 1-SRI002. This program was also supported by the Australian Research Council (ARC) through grant DP180102045. The views expressed herein are not necessarily the views of the Australian Government, and the Australian Government does not accept responsibility for any information or advice contained here. The valuable feedback on the original draft by the anonymous reviewers is gratefully acknowledged.
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Chisti, M.Y., Moo-Young, M., 1987. Airlift reactors: Characteristics, applications and design considerations. Chem. Eng. Commun. 60, 195–242. Chisti, M.Y., Moo-Young, M., 1988. Gas holdup in pneumatic reactors. Chem. Eng. J. 38, 149–152. Chisti, Y., Moo-Young, M., 1993. Improve the performance of airlift reactors. Chem. Eng. Prog. 89(6), 38–45. Chisti, M.Y., Halard, B., Moo-Young, M., 1988. Liquid circulation in airlift reactors. Chem. Eng. Sci. 43, 451–457. Deckwer, W.-D., 1992. Bubble Column Reactors. Wiley, Chichester. Hikita, H., Kikukawa, H., 1973. Gas hold-up in bubble columns: effect of liquid properties. Bull. Univ. Osaka Prefecture Ser. A., Engng. Nat. Sci., 22(2), 151–160. Hikita, H., Asai, S., Tanigawa, K., Segawa, K., Kitao, M., 1980. Gas hold-up in bubble columns. Chem. Eng. J. 20, 59–67. Jafarian, M., Arjomandi, M., Abdollahi, M.R., Nathan, G.J., 2017a. Concentrated solar receiver and reactor systems comprising heat transfer fluid. International Patent Application No. PCT/AU2018/050034
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Merchuk, J.C., 2003. Airlift bioreactors: Review of recent advances. Can. J. Chem. Eng. 81, 324–337. Merchuk, J.C., Siegel, M.H., 1988. Air-lift reactors in chemical and biological technology. J. Chem. Technol. Biotechnol. 41, 105–120. Moo-Young, M., Blanch, H.W., 1981. Design of biochemical reactors: mass transfer criteria for simple and complex systems. Adv. Biochem. Eng. 19, 1–69. Sarafraz, M.M., Jafarian, M., Arjomandi, M., Nathan, G.J., 2017a. Potential use of liquid metal oxides for chemical looping gasification: A thermodynamic assessment. Appl. Energy 195, 702–712. Sarafraz, M.M., Jafarian, M., Arjomandi, M., Nathan, G.J., 2017b. The relative performance of alternative oxygen carriers for liquid chemical looping combustion and gasification. Int. J. Hydrogen Energy 42, 16396–16407. Sarafraz, M.M., Jafarian, M., Arjomandi, M., Nathan, G.J., 2018. Potential of molten lead oxide for liquid chemical looping gasification (LCLG): A thermochemical analysis. Int. J. Hydrogen Energy 43, 4195–4210. Sarafraz, M.M., Jafarian, M., Arjomandi, M., Nathan, G.J., 2019a. Experimental investigation of the reduction of liquid bismuth oxide with graphite. Fuel Process. Technol. 188, 110– 117. Sarafraz, M.M., Jafarian, M., Arjomandi, M., Nathan, G.J., 2019b. The thermo-chemical potential liquid chemical looping gasification with bismuth oxide. Int. J. Hydrogen Energy 44, 8038–8050. Sasaki, S., Uchida, K., Hayashi, K., Tomiyama, A., 2017. Effects of column diameter and liquid height on gas holdup in air-water bubble columns. Exp. Therm. Fluid Sci. 82, 359–366.
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Upham, D.C., Snodgrass, Z.R., Zavareh, M.T., McConnaughy, T.B., Gordon, M.J., Metiu, H., McFarland, E.W., 2017. Molten salt chemical looping for reactive separation of HBr in a halogen-based natural gas conversion process. Chem. Eng. Sci. 160, 245–253. Zhang, T., Wang, T., Wang, J., 2005. Mathematical modeling of the residence time distribution in loop reactors. Chem. Eng. Proc. 44, 1221–1227. Zhang, T., We, C., Ren, Y., Feng, C., Wu, H., 2017. Advances in airlift reactors: modified design and optimization of operation conditions. Rev. Chem. Eng. 33, 163–182.
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Table 1 Summary of the experimental conditions used.
Gas-free liquid height
Sparger nozzle diameter
Superficial gas velocity (Us, m s1)
(HL, mm)b
(d, mm)b
Column 1
Column 2
1c
300
1.4, 2.0, 2.8, 4.0 and 5.6
0.01 to 0.16
0.01 to 0.16
2
300
2.8
0.039
0 to 0.16
3c
212, 300, 424 and 600
2.8
0.01 to 0.16
0.01 to 0.16
Experimental seta
a In
all cases, the diameters of the columns 1 and 2 were identical at 85 mm.
b These c
variables had identical values in the two bubble columns.
For experimental sets 1 and 3, the superficial gas velocity in the two columns was kept equal for all combinations of sparger nozzle diameters
and heights of the gas-free liquid.
26
Figure captions Fig. 1. Gas-lift reactor with dual interconnected isothermal bubble columns (Jafarian et al., 2017a; Jafarian et al., 2017b). The two bubble columns functioned as riser zones of the circulation loop. Both the risers were injected with a gas. The gas-liquid dispersion leaving the upper zones of the bubble columns was further degassed in gas-liquid separators (degassers) and the remaining liquid was returned to the bottom of the adjacent column. The gas sparged in the bubble columns provided the gas-lift action that moved the liquid between the columns.
Fig. 2. Gas holdup (G) versus superficial gas velocity (Us) in column 1 (a) and column 2 (b). The data are for the experimental set 1 (Table 1). The best fit empirical correlations for gas holdup in the two columns are shown. Identical nozzles of hole diameter d were used in both columns.
Fig. 3. Dependence of the circulation flow rate (Qc) between the bubble columns and the mean residence time (r) of water in each bubble column, on superficial velocity (Us) of gas in the columns. (Us values were identical in the two columns at any time.) Identical nozzles of diameter d were used for gas injection in the two columns. The conditions shown are for experimental set 1 (Table 1). The equation for Qc is based on measured data. The residence time (r) was calculated by substituting the above shown empirical correlation for Qc in Eq. (15).
Fig. 4. Comparison of Eq. (17) (black line) with published correlations (Chisti and MooYoung, 1988; Chisti, 1989; Chisti, 1998) for bubble-flow (red line) and churn-turbulent flow (blue line) regimes for air–water in conventionally operated bubble columns of diverse sizes.
27
Fig. 5. Energy dissipation rate (EF) due to circulation of water between the bubble columns versus superficial gas velocity (Us). The gas in both columns was injected through identical nozzles with a hole diameter d and the Us values were identical. The conditions shown are for experimental set 1 (Table 1). The best fit empirical correlation for the entire data set is shown. EF values were calculated using Eq. (10).
Fig. 6. Dependence of the fractional power input attributable to gas expansion () within the columns, on the superficial gas velocity (Us) in the columns. The gas in both columns was injected through nozzles having a hole diameter d and the Us values were identical in the two columns. The conditions shown are for experimental set 1 (Table 1). values were calculated using Eq. (11).
Fig. 7. Comparison of measured induced liquid circulation rate (Qc) for the modes of operation with: (a) both columns having equal values superficial gas velocities (Us,C1 = Us,C2) at any instance; and (b) column 1 having a fixed aeration rate Us,C1 = 0.039 m s1 for different values of the aeration velocity Us,C2 in column 2. The sparger hole diameter d in both columns was always 2.8 mm and the unaerated liquid height HL was always 300 mm.
Fig. 8. Dependence of the circulation flow rate (Qc) between the bubble columns and the mean residence time (r) of water in each bubble column, on superficial velocity (Us,C2) of the gas injected in column 2. Superficial air velocity (Us,C1) in column 1 remained fixed at 0.039 m s1. Data are shown are for conditions of experimental set 2 (Table 1). Best fit empirical correlations are shown.
28
Fig. 9. Gas holdup (G,C1, G,C2; subscripts C1 and C2 denote columns 1 and 2, respectively) values in the two columns versus superficial gas velocity (Us,C2) in column 2. The superficial gas velocity in column 1 was kept constant at 0.039 m s1. The data are for the conditions in experimental set 2 (Table 1). The best fit empirical equation for gas holdup in column 2 is shown.
Fig. 10. Comparison of the energy dissipation rate (EF) due to circulation of water between the columns, for modes of operation with: (a) column 1 having a fixed aeration rate Us,C1 = 0.039 m s1 for different values of the aeration velocity Us,C2 in column 2; and (b) both columns having equal values superficial gas velocities (Us,C1 = Us,C2) at any instance. The best fit empirical equation for the operation mode (a) is shown. EF values were calculated using Eq. (10). The sparger hole diameter d in both columns was always 2.8 mm and the unaerated liquid height HL was always 300 mm.
Fig. 11. Dependence of the circulation flow rate (Qc) between the bubble columns on superficial gas velocity (Us, identical in both columns) in the columns. Data are shown for identical specified heights (HL) of gas-free water in the two columns (experimental set 3; Table 1). In all case the sparger hole diameter was the same (d = 2.8 mm). Best fit empirical correlations for the data are shown.
Fig. 12. Comparison of the experimentally measured Qc values (Fig. 11) with predictions of Eq. (23). Most of the measured data were within 15% of the solid line, i.e. the line of exact agreement (y = x).
29
Fig. 13. Dependence of the mean residence time (r) of water in each bubble column, on superficial gas velocity (Us) and gas-free liquid height (HL) in the two columns. Both columns had the same values of Us and HL. Data shown are for conditions of experimental set 3 (Table 1).
Fig. 14. Dependence of the gas holdup on superficial gas velocity in bubble column 1 (a) and bubble column 2 (b). Data are shown for various heights (HL; identical in both columns) of gas-free water in the two columns. Superficial gas velocity (Us) was the same in both columns. Data shown are for conditions of experimental set 3 (Table 1). Best fit empirical equations describing the data are shown.
Fig. 15. Dependence of the fractional power input attributable to gas expansion () within the columns, on the superficial gas velocity (Us; identical in both columns) in the columns. Data are shown for various heights (HL; identical in both columns) of gas-free water in the two columns. Data shown are for the conditions of experimental set 3 (Table 1). -values were calculated using Eq. (11). All columns had the same diameter (d = 2.8 mm) of aeration nozzle.
30
Fig. 1
31
Fig. 2
32
Fig. 3
33
G = 49Us0.46 G = 247Us0.97
G (%)
10
G = 63.33Us0.80
Present work (Equation 17) Conventional bubble columns (bubble-flow regime) Conventional bubble columns (churn-turblent flow)
1 0.01
0.1 1
Us (m s )
Fig. 4
34
Fig. 5
35
Fig. 6
36
Fig. 7
37
2.5
110
Qc = 1.02 + 3.06Us,C20.564
100 90
1.5
3
80
r = 105.13 113.39Us,C20.385
1.0
70 60
0.5
Qc
50
r 0.0 0.00
0.02
0.04
0.06
0.08
0.10 1
Us,C2 (m s )
Fig. 8
38
0.12
0.14
0.16
40 0.18
r (s)
1
Qc (dm min )
2.0
14
G,C1, G,C2 (%)
12
Column 1 (Us,C1 = 0.039 m s1) Column 2
10
G,C2 = 67.57Us,C20.841
8 6 4
G,C1
2 0 0.00
0.02
0.04
0.06
0.08
0.10 1
Us,C2 (m s )
Fig. 9
39
0.12
0.14
0.16
0.18
0.05
EF (W)
0.04
(b) Us,C1 = Us,C2
0.03
(a) 0.02
0.01
EF = 2.3x103 + 0.145Us,C2 + 0.121Us,C22 0.00 0.00
0.02
0.04
0.06
0.08
0.10
0.12
Us,C2 (m s1)
Fig. 10
40
0.14
0.16
0.18
0.20
10
Qc = 10.43Us
0.457
3
0.534
Qc = 7.03Us
1
Qc (dm min )
Qc = 9.84Us
1
0.501
Qc = 4.70Us0.469 HL (mm) 600 424 300 212
0.1 0.01
0.1
1 1
Us (m s )
Fig. 11
41
5
4
15%
3
1
Predicted Qc (dm min )
+15%
3
2
HL (mm)
212 300 424 600
1
0
0
1
2
3
4 3
1
Experimental Qc (dm min )
Fig. 12
42
5
Fig. 13
43
Fig. 14
44
Fig. 15
45
Highlights A novel two-column gas-lift reactor configuration was developed. Reactor comprised two interconnected riser zones forming an -shaped loop. Mixing and circulation were achieved solely by a gas-lift action. Possible applications in chemical looping with different gases in the two columns. Correlations were developed for gas holdup and liquid circulation rate.
46
CRediT authorship contribution statement Mehdi Jafarian: Conceptualization, Investigation, Methodology, Data Curation, Formal analysis, Validation, Visualization, Writing - Original Draft. Yusuf Chisti: Investigation, Visualization, Formal analysis, Writing-review& editing. Graham J. Nathan: Conceptualization, Supervision, Formal analysis, Methodology, Writing review & editing, Supervision, Funding acquisition.
47