Mass transfer characteristic of ammonia escape and energy penalty analysis in the regeneration process

Applied Energy 257 (xxxx) xxxx

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Mass transfer characteristic of ammonia escape and energy penalty analysis in the regeneration process ⁎

Fengming Chua, Qianhong Gaoa, Shang Lib, Guoan Yanga, , Yan Luoa, a b

⁎

College of Mechanical and Electrical Engineering, Beijing University of Chemical Technology, Beijing 100029, China China Academy of Information and Communication Technology, Beijing 100191, China

H I GH L IG H T S

computational model of ammonia regeneration model in packed column is built. • AThenovel overall mass transfer coefficient of ammonia escape is defined. • Increasing CO loading can inhibit ammonia escape and reduce energy consumption. • The accurateof three parts of regeneration energy consumption is gained. • Energy penalty is influenced by the initial ammonia concentration and CO loading. • 2

2

A R T I C LE I N FO

A B S T R A C T

Keywords: Carbon dioxide capture Ammonia escape Stripper Mass transfer Regeneration

The energy penalty and ammonia escape of the regeneration process are the main barriers for the ammonia technology application, so it is of significance for the commercial application to investigate the dominating parameters influencing the ammonia escape and energy consumption performance. A rigorous computational model of the ammonia regeneration process is established, based on which the ammonia regeneration and escape processes can be predicted accurately. The ammonia escape and energy penalty of the regeneration process are studied and the results show that the ammonia escape performance is controlled by the liquid film. The increasing of the liquid inlet temperature, initial ammonia concentration and CO2 loading can reduce the regeneration energy consumption obviously. Especially, the energy consumption of 343.15 K liquid inlet temperature is almost three times to that of 368.15 K liquid inlet temperature. This work can contribute to the ammonia escape inhibition and low energy consumption in the ammonia regeneration process for the industrial application of the ammonia technology.

1. Introduction CO2 capture and storage (CCS) is regarded as the most promising technology to mitigate climate change, which is due to the serious CO2 emissions [1]. Various technologies were proposed to reduce the CO2 emission, including the membrane separation [2], physical adsorption [3], chemical absorption, cryogenic methods [4], photocatalytic reduction, convert to methanol [5]. Among these methods, the postcombustion technology is regarded as the promising commercial application, which is based on the chemical absorption [6]. In the development context of the post-combustion CO2 capture technology, the monoethanolamine (MEA) solution and the ammonia solution are expected to play the major roles in the absorbents [7]. Compared with MEA solution, the aqueous ammonia solution has much higher loading

⁎

capacity and requires lower energy for regeneration [8]. However, the ammonia technology has two less favorable views of the application, which are the regeneration energy consumption and the ammonia escape. Due to the high temperature in the stripper, the ammonia escape is much more serious than that in the absorber. Therefore, it is of benefit to study the ammonia escape process and the energy consumption in the stripper. In the past several decades, numerous researches involving the ammonia regeneration process have been carried out. Ma et al. [9] measured the activation energy and reaction rate constant of the ammonia regeneration reaction, and declared that the desorption reaction was a second order reaction. Ma et al. [10] studied the ammonia regeneration process by the experimental method, and the results showed that the CO2 desorption percentage increased with the rising of the

Corresponding authors. E-mail addresses: [email protected] (G. Yang), [email protected] (Y. Luo).

https://doi.org/10.1016/j.apenergy.2019.113975 Received 4 July 2019; Received in revised form 14 September 2019; Accepted 8 October 2019 0306-2619/ © 2019 Elsevier Ltd. All rights reserved.

Please cite this article as: Fengming Chu, et al., Applied Energy, https://doi.org/10.1016/j.apenergy.2019.113975

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Nomenclature

p pH2O,G pH2O pNH3 pNH3,i SC

surface area per unit volume of packed bed, m2/m3 contact area of the gas and liquid phases in the stripper, m2/m3 Ci mass fraction of NH4HCO3 in the rich solution, dimensionless cP_G specific heat of gas phase, J/(kg·K) cp,r specific heat capacity of the rich solution, kJ/(kg·K) CNH3 mass fraction of ammonia in the liquid phase, dimensionless CNH3,G mass fraction of ammonia in the gas phase, dimensionless Cw mass fraction of water in the liquid phase, dimensionless CH2O,G mass fraction of water in the gas phase, dimensionless DL,i molecular diffusivity of NH4HCO3 in the rich solution, m2/ s DL,NH3 molecular diffusivity of NH3 in the rich solution, m2/s Dt,i turbulent diffusivity of NH4HCO3 in the rich solution, m2/ s Dt,NH3 molecular diffusivity of NH3 in the rich solution, m2/s FLG interface drags force between gas phase and liquid phase, N/m3 FLS flow resistance created by the packings, N/m3 g gravitational acceleration, m/s2 GM gas phase mass velocity, kg/m2/h G gas phase molar flow rate, kmol/m2/s h liquid phase volume fraction based on pore space, dimensionless hexc convective heat transfer coefficient, w/m2/K hg volume fraction of gas phase based on pore space, dimensionless HNH3 Henry’s constant of ammonia, kPa·m3/kmol k turbulent kinetic energy, m2/s2 KG,H2O overall mass transfer coefficient of H2O, kmol/kPa/m2/s kG,NH3 gas phase mass transfer coefficients of NH3, kmol/m2/s/ kPa kL,NH3 liquid phase mass transfer coefficients of NH3, m/s KG,NH3aw over all mass transfer coefficients of NH3 escape, kmol/ m3/s/kPa MH2O molecular weight of H2O, kg/kmol MCO2 molecular weight of CO2, kg/kmol MNH3 molecular weight of NH3, kg/kmol MNH4HCO3 molecular weight of NH4HCO3, kg/kmol aT aw

SCO2 SH2O Sm ST_G ST_L SNH3 SNH3_G Sw TG XNH3 XNH3,i yCO2 yNH3 yNH3,i U Ug V

pressure, kPa partial pressures of H2O in the main gas phase, kPa saturated vapor pressure of the water, kPa NH3 partial pressures in the main gas phase, kPa NH3 partial pressures at the interface, kPa source term of NH4HCO3 mass fraction equation, kg/ (m3·s) source term of the CO2 mass fraction equation, kg/(m3·s) source term of H2O mass fraction in the gas phase, kg/ (m3·s) source term of the liquid phase continuity equation, kg/ (m3·s) source term of the gas phase energy equation, J/(m3·s) source term of the liquid phase energy equation, J/(m3·s) source term of NH3 mass fraction equation in the liquid phase, kg/(m3·s) source term of NH3 mass fraction in the gas phase, kg/ (m3·s) source term of water mass fraction equation in the liquid phase, kg/(m3·s) temperature of the gas phase, K concentrations of ammonia in the main liquid phase, kmol/m3 concentrations of ammonia at the interface, kmol/m3 CO2 volume fraction in gas phase, dimensionless. NH3 volume fraction in gas phase, dimensionless. NH3 volume fraction at the interface, dimensionless. liquid phase interstitial velocity vector in the x direction, m/s velocity of the gas phase, m/s liquid phase interstitial velocity vector in the r direction, m/s

Greek symbols αL,αt ε μ,μt,μeff μg ρ ρg

molecular and turbulent thermal diffusivities, m2/s turbulent dissipation rate, m2/s3 molecular, turbulent and effective viscosities of liquid phase, kg/(m·s) gas phase viscosity, Pa·s liquid density, kg/m3 gas phase density, kg/m3

regeneration energy can be reduced by the addition of piperazine. Zhai et al. [18] proposed three integration methods of the aqueous ammonia based CO2 capture technology integration with power plants. Zhang and Guo [19] studied the effects of the stripper size on the ammonia regeneration process by a rate-based model. Besides the regeneration energy consumption, the ammonia loss is also the main barrier to the application of the ammonia technology. Wang et al. [20] suggested that the chilled ammonia process can inhibit the ammonia evaporation effectively. Zhang and Guo [21] investigated an ammonia abatement system, in which the lean NH3 solvent was used to absorb the escaping NH3. Asif et al. [22] believed the vacuum membrane distillation can solve the ammonia loss problem by recovering ammonia from ammonia wash water. The similar phenomenon can be found in the literature of Fang et al. [23]. Li et al. [24] pointed out that ammonia recovery process was feasible and promising prospect toward industrial application. Wang et al. [25] described the mass transfer process of the ammonia escape and pointed out that the ammonia escape rate was affected by the ammonia concentration, gas and liquid flow rates, reaction temperature and desorption pressure qualitatively.

temperature, solution concentration and CO2 loading. Yeh et al. [11] pointed out that ammonium bicarbonate required the less regeneration energy than any other ammonium compounds. Yu and Wang [12] built a rate-based model of the ammonia regeneration process, based on which the regeneration energy consumption was analyzed. Ullah et al. [13] investigated the impacts of a novel capacitive deionization device on the ammonia regeneration process by the rate-based model, and pointed out that the regeneration energy consumption can be reduced by 37.5%. Zhang and Guo [14] established a comprehensive model for the ammonia regeneration process, which took hydrodynamics, thermodynamic, rate-based mass transfer and reaction into consideration simultaneously on Aspen Plus. Zhang and Guo [15] conducted a reaction sensitivity analysis of the ammonia regeneration process by the rate-based model, and reported that the bicarbonate decomposition reaction played an important roles on the ammonia regeneration process. Zhang and Guo [16] proposed a capacitive ion separation (CIS) device, which can reduce the required energy of ammonia regeneration. The regeneration energy can be reduced by up to 35% by the CIS device. Lu et al. [17] developed a rate-based model for ammonia regeneration process using Aspen Plus, reporting the ammonia 2

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2NH 4 HCO3 → (NH 4 )2 CO3 (aq) + CO2 (g) + H2 O(l) H2 = 27.51 kJ·mol−1

The aforementioned researching emphases were placed on the energy penalty of the regeneration process. However, the ammonia escape in the stripper was rarely reported. The heat and mass transfer process of the regeneration process is very complex, which contains that CO2 desorbing, H2O evaporation, NH3 escape, chemical absorption heat, convective heat transfer and the reaction. Therefore, it is difficult to investigate the ammonia regeneration process only through the experimental method. In order to effectively explore the ammonia regeneration process it is needed to establish a numerical model should, which can take all the mass transfer process, hydrodynamics, thermodynamic, and regeneration reaction into consideration. A rigorous ammonia regeneration simulation model was established in this paper, based on which the ammonia escape mass transfer performance and energy consumption were analyzed in details. The influences of the main working conditions on the ammonia escape and energy penalty were clarified and the dominating parameters for the ammonia escape were reported, which were both making the ammonia technology more viable.

(2)

NH 4 HCO3 (aq) → CO2 (aq) + NH3 (aq) + H2 O(l)

H3 = 64.87

kJ·mol−1 (3)

(NH 4 )2 CO3 (aq) → CO2 (aq) + 2NH3 (aq) + H2 O(l)

H4 = 102.23 kJ·mol−1 (4)

mol−1

(5)

kJ·mol−1

(6)

CO2 (aq) → CO2 (g) H5 = 19.75 kJ

NH3 (aq) → NH3 (g)

H2 O(l) → H2 O(g)

H6 = 34.35

H7 = 44.00 kJ·mol−1

(7)

2. Models

where H1, H2, H3 and H4 are the reaction heat, respectively. H5, H6 and H7 represent the physical heat of the evaporation. Eqs. (1)–(4) are the hydrolysis process of the bicarbonate, carbonate and carbamate in the carbonization ammonia solution, which are endothermic reactions. However, it is of difficulty to investigate all the reaction processes by the numerical method. Therefore, the main absorbing product is assumed to be the ammonium bicarbonate (NH4HCO3). Therefore, the regeneration reaction is presented as follows:

2.1. Reaction mechanism

NH 4+ (aq) + HCO3- (aq) → CO2 (aq) + NH3 (aq) + H2 O(l)

As shown in Fig. 1, the regeneration reaction occurs in the liquid solution, which is the reversed reaction of absorption. CO2 is desorbed from the rich solution and then goes into the gas phase. Owing to the complex species produced in the absorbing process, there are the complex species equilibrium in rich solution. The main reactions and mass transfer of the ammonia regeneration process are presented as follows [9–11]:

NH2 COONH 4 (aq) + H2 O(l) → NH 4 HCO3 (aq) + NH3 (aq) mol−1

(8)

2.2. Modeling approach and assumptions The CFD model in this paper is mainly based on the Representative Elementary Volume (REV) model. The flow in the stripper is assumed to be steady, which is due to the pseudo-steady-state theory [26]. The liquid phase flow is regarded as the pseudo-continuous flow and the density is assumed to be constant [27–29]. Besides, the following assumptions should also be made:

H1 = 30.70 kJ· (1) The flow in the stripper is axis-symmetric.

(1)

Fig. 1. Schematics of CCS system and the species transfer in the stripper. 3

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(2) The regeneration reaction is assumed to be steady. (3) There is no heat exchange between the stripper and the environment. (4) The reaction heat of the endothermic desorption process is only provided by the liquid phase instantaneously. (5) The change of the gas phase mass is not big enough to change the gas velocity and density. (6) Because the amount of CO2 in the gas phase is low, the velocity and density of the gas phase are assumed to be constant.

μt ⎞ ∂ε ⎞ μ ∂ε ∂ (hρUε ) ∂ ⎡ ⎛ 1 ∂ (hρrVε ) 1 ∂ ⎛ ⎛ + − h μ + t⎞ ⎤ − ⎜hr μ + ⎟ ⎥ ∂x ∂x ⎢ σ ∂ σ x r ∂ r r r ε ε ⎠ ∂r ⎠ ⎝ ⎠ ⎝ ⎣ ⎦ ⎝ ⎜

⎟

⎜

⎟

ε2 ε V 2 ∂U ⎞2 ⎛ ∂V ⎞2 ∂U ∂V ⎞2⎤ 2 ⎜⎛ ⎛ = c1·hμt ·⎡ + + ⎛ ⎞ ⎟⎞ + ⎛ + − c2·ρh ⎢ ⎥ k k ⎣ ⎝ ⎝ ∂x ⎠ ∂x ⎠ ⎦ ⎝ r ⎠ ⎠ ⎝ ∂r ⎝ ∂r ⎠ (16) The constants in the turbulent model are set as Cμ = 0.09, σk = 1.0, σε = 1.3, c1 = 1.44 and c2 = 1.92 [27–29]. 2.3.2. REV-heat transfer model

2.3. Governing equations

(1) Rich solution energy equation

2.3.1. REV-fluid flow model

∂ (ρhcp, r UT )

1 ∂ (ρhcp, r rVT ) + r ∂r ∂x ∂ ⎡ ∂T 1 ∂ ⎡ ∂T = ρhcp, r (αL + αt ) ⎤ + ρhcp, r (αL + αt ) r ⎤ + ST _L ∂x ⎣ ∂x ⎦ r ∂r ⎣ ∂r ⎦

(1) Rich solution continuity equation

∂ (ρhU ) 1 ∂ (ρhrV ) = Sm + r ∂r ∂x

(9)

(17)

where ρ (kg/m3) represents the rich solution density in the stripper. h is the volume fraction of the rich solution. U, V (m/s) are the velocity vectors in the x and r directions of the rich solution in the stripper. Sm (kg/m3/s) is the source term of the continuity equation, which is the mass transferred from the rich solution to gas phase per unit volume and time.

where ST_L(J/m3/s) is the source term of the rich solution energy equation, which is the sum of the absorbing heat of the endothermicregeneration reaction and the physical heat transfer with the gas phase. cp,r,(kJ/kg/K) and T(K) represent the specific heat capacity and temperature of liquid phase. αL and αt (m2/s) are the molecular and turbulent thermal diffusivities of liquid phase. The impacts of the turbulent flow the heat transfer performance are modelled by the turbulent thermal diffusivities αt, which is used the analogy method with the REV-fluid flow model.

(2) Rich solution momentum equation The x direction:

1

∂ (ρhUU ) 1 ∂ (ρhrUV ) 2 1 ∂ ∂ ⎡ ∂U hμ ⎛2 − − (∇ ·U) ⎞ ⎤ = + ⎢ eff ⎝ ∂x ⎥ r 3 r ∂r ∂r ∂x ⎣ ∂x ⎠⎦ p ∂ U V ∂ ∂ ⎡rhμ ⎛ ⎞⎤ − h + h (ρg + FLS, x ) + FLG + ⎢ eff ⎝ ∂r ⎥ ∂x ∂x ⎠ ⎦ ⎣

k t2 ⎞ αt = ct 0 k ⎛⎜ ⎟ ⎝ ε εt ⎠ (10)

εt = αL

∂U 1 ∂ (rV ) + ∂x r ∂r

μeff = μ + μt k2 ε

(12)

(2) Temperature variance t 2 equation

(13)

∂ (ρhU t 2) 1 ∂ (rρhV t 2) + r ∂r ∂x =

where g (m/s ) and p (kPa) are the gravitational acceleration and the pressure, respectively. FLG (N/m) and FLS (N/m) are the drag force with the gas phase and the body force created by the packing materials. μ, μt and μeff represent the molecular, turbulent and effective viscosities of the liquid phase, respectively. k(m2/s2) and ε(m2/s3) are the turbulent kinetic energy and turbulent dissipation rate, which is used to determine the turbulent viscosity μt.

μt ⎞ ∂k ⎞ μt ⎞ ∂k ⎞ ∂ (ρhUk ) ∂ ⎛ ⎛ 1 ∂ (hρrVk ) 1 ∂ ⎛ ⎛ − + ⎜h μ + ⎟ − ⎜rh μ + ⎟ ∂r ∂x ⎝ ⎝ σk ⎠ ∂r ⎠ ∂x σk ⎠ ∂x ⎠ r ∂r ⎝ ⎝ f r 2

2

⎟

2

V ∂U ∂V ⎞ ⎤ ⎛ ⎛ ∂U ⎞ ⎛ ∂V ⎞ ⎛ ⎞⎞ ⎛ = hμt ⎡ ⎢2 ⎜ ⎝ ∂x ⎠ + ⎝ ∂r ⎠ + ⎝ r ⎠ ⎟ + ⎝ ∂r + ∂x ⎠ ⎥ − hρε ⎠ ⎣ ⎝ ⎦

⎜

⎟

⎜

⎟

(21)

∂ (ρhUεt ) 1 ∂ (ρhrVεt ) + r ∂r ∂x =

⎜

1 ∂ ⎡ α ∂t2 ⎤ ∂t2 ⎤ ∂ ⎡ ⎛ αt ρh ρhr ⎛ t + αL ⎞ − 2ρhαt + + αL ⎞ ⎢ ⎥ σ x r r σ ∂ ∂ ∂x ⎢ t t ⎝ ⎠ ∂r ⎥ ⎠ ⎦ ⎣ ⎦ ⎣ ⎝ ⎛ ∂T ∂T + ∂T ∂T ⎞ − 2ρhεt ∂r ∂r ⎠ ⎝ ∂x ∂x

(3) Temperature variance dissipation rate εt equation

(3) Turbulent kinetic energy equation ⎟

(20)

(11)

2

⎜

∂t ∂t ∂x j ∂x j

The temperature variance and temperature variance dissipation rate of the rich solution are defined by the analogy with the turbulent kinetic energy equation and turbulent dissipation rate equation.

(14)

2

(19)

t 2 = tt

∂ (ρhUV ) 1 ∂ (ρhrVV ) 1 ∂ ⎡ 2 ∂V ∂ rhμ ⎛2 − − (∇ ·U) ⎞ ⎤ = + ⎢ eff ⎝ ∂r ⎥ 3 r r ∂r ⎣ ∂r ∂x ∂x ⎠⎦ 2hμeff ⎡hμ ⎛ ∂U + ∂V ⎞ ⎤ − h ∂P − 2hμ V + (∇ ·U) + hFLS, r eff 2 ⎢ eff ⎝ ∂r ⎥ 3 r ∂x ⎠ ⎦ ∂r ⎣

μt = ρCμ

(18)

where t 2 and εt (1/s) are the temperature variance and temperature variance dissipation rate, which is obtained as follows:

The r direction:

∇ ·U=

2

1 ∂ ⎡ α ∂ε ∂ε ∂ ⎡ ⎛ αt ρh ρhr ⎛ t + αL ⎞ t ⎤ − Ct1 + αL ⎞ t ⎤ + ⎥ ⎢ σ x r r σ ∂r ⎥ ∂ ∂ ∂x ⎢ t t ⎝ ⎠ ⎝ ⎠ ⎦ ⎦ ⎣ ⎣ 2 ε ∂ T ∂ T ∂ T ∂ T ε εε t t t ⎛ ⎞ ρh + − Ct 2 2 − Ct 3 ∂r ∂r ⎠ t 2 k t ⎝ ∂x ∂x ⎜

⎟

⎜

t2

⎟

(22)

The parameters in the − εt model are given as Ct0 = 0.11, Ct1 = 1.8, Ct2 = 2.2, Ct3 = 0.8, σt = 1.0 and σεt = 1.0 [27–29].

(15)

(4) Gas energy equation

(4) Turbulent dissipation rate equation

4

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∂ (ρg hg cp _G Ug TG ) ∂x

= ST _G

∂ (ρhUCNH3 ) 1 ∂ (ρhrVCNH3 ) + ∂x r ∂r ∂CNH3 ⎤ ∂ ⎡ 1 ∂ = ρh (DL,NH3 + Dt ,NH3 ) + ⎥ ∂x ⎢ ∂ x r ∂r ⎣ ⎦

(23)

where ST_G(J/m3/s) is the energy equation source term of gas phase, which is the heat transferred with liquid phase. cP_G (J/(kg·K)) and ρg (kg/m3) represent the specific heat and density of the gas phase, respectively. Ug (m/s) and hg are the gas phase velocity and volume fraction in the stripper.

⎡ρh (DL,NH + Dt ,NH ) r ∂CNH3 ⎤ + SNH 3 3 3 ⎢ ∂r ⎥ ⎣ ⎦

2.3.3. REV-mass transfer model (1) NH4HCO3 mass fraction equation

∂ (ρhUCi ) 1 ∂ (ρhrVCi ) + r ∂r ∂x ∂ ⎡ ∂C 1 ∂ ⎡ ∂C = ρh (DL, i + Dt , i ) i ⎤ + ρh (DL, i + Dt , i ) r i ⎤ + SC ∂x ⎣ ∂x ⎦ r ∂r ⎣ ∂r ⎦

(24)

(5) Water mass fraction equation in the liquid phase

where Ci is the mass fraction of NH4HCO3 in the rich solution. Sc (kg/ m3/s) is the source term of NH4HCO3 mass fraction equation, which is the consuming mass per unit volume and time. DL,i (m2/s) and Dt,i (m2/ s) are the molecular and turbulent diffusivities for mass transfer of NH4HCO3 in the liquid phase, respectively. The impacts of the turbulent flow the mass transfer performance are modelled by the turbulent diffusivities Dt,i, which is used the analogy method with the REV-fluid flow model and simplified as follows: 1

k ci2 ⎤ Dt , i = Cc0 k ⎡ ⎢ ε εC ⎥ i⎦ ⎣

(31) where Cw represents the mass fraction of water. Sw (kg/m3/s) is the source term of water mass fraction equation, which represents the mass transfer of water from liquid phase to gas phase per unit volume and time. DL,w (m2/s) and Dt,w (m2/s) are the molecular and turbulent diffusivities for mass transfer of water, respectively. Dt,w can be obtained according to the above part.

(25)

The concentration variance

ci2

∂ (ρhUCw ) 1 ∂ (ρhrVCw ) + ∂x r ∂r ∂ ⎡ ∂C 1 ∂ ⎡ ∂C = ρh (DL, w + Dt , w ) w ⎤ + ρh (DL, w + Dt , w ) r w ⎤ + Sw ∂x ⎣ ∂x ⎦ r ∂r ⎣ ∂r ⎦

2

ci2

is expressed as:

= ci ci

(6) CO2 mass fraction equation in the gas phase CCO2

(26)

∂ (ρg hg Ug CCO2 )

The concentration variance dissipation rate εci (1/s) is defined by:

∂c ∂c εci = DL, i i i ∂x j ∂x j

∂x

(32)

where SCO2 (kg/m /s) represents the source term of CO2 mass fraction owing to the desorption of rich solution.

(27)

(7) NH3 mass fraction in the gas phase CNH3,G

∂ (ρg hg Ug CNH3, G ) ∂x

(33)

where SNH3_G (kg/m /s) represents the source term of NH3 mass fraction equation due to the NH3 escape.

∂ (ρhUci2 ) 1 ∂ (ρhrVci2 ) + r ∂x ∂r

(8) Water mass fraction equation in the gas phase

∂ (ρg hg Ug CH2O, G )

Dt , i ⎞ ∂ci2 ⎤ Dt , i ⎞ ∂ci2 ⎤ 1 ∂ ⎡ ⎛ ∂ ⎡ ⎛ ρh DL, i + ρh DL, i + − 2ρhDt , i + ⎥ ⎢ ⎢ σc ⎠ ∂r ⎥ σc ⎠ ∂x ⎦ r ∂r ⎣ ⎝ ∂x ⎣ ⎝ ⎦ ⎟

= SNH3_G 3

(2) Concentration variance ci 2 equation

⎜

= SCO2 3

The concentration variance and concentration variance dissipation rate are determined by the analogy with the turbulent kinetic energy equation and turbulent dissipation rate equation, which is used to modify the mass transfer performance under the turbulent flow condition.

=

(30)

where CNH3 is the mass fraction of NH3 in the liquid phase. SNH3 (kg/ m3/s) is the source term of mass conservation equation, which is the difference between the regeneration reaction production and ammonia escape mass per unit volume and time. DL,NH3 (m2/s) and Dt,NH3 (m2/s) are the molecular and the turbulent diffusivities of NH3 in the rich solution. Dt,NH3 can be obtained according to the above part.

⎜

∂x

⎟

⎛ ∂Ci ∂Ci + ∂Ci ∂Ci ⎞ − 2ρhεCi ∂r ∂r ⎠ ⎝ ∂x ∂x

= SH2O

(34)

3

where SH2O (kg/m /s) represents the source term of H2O mass fraction in the gas phase owing to the evaporation.

(28)

(3) Concentration variance dissipation rate εci equation

2.4. Correlative terms

∂ (ρhUεCi ) 1 ∂ (ρhVεCi ) + ∂x ∂r r

2.4.1. Source term of water evaporation The source term of H2O conservation equation originates from the evaporation of water, which can be calculated as follows [30,31]:

=

Dt , i ⎞ ∂εCi ⎤ Dt , i ⎞ ∂εCi ⎤ ∂ ⎡ ⎛ 1 ∂ ⎡ ⎛ ρh ⎜DL, i + ⎟ ⎥ + r ∂r ⎢ρhr DL, i + σc ∂r ⎥ − Cc1 Dt , i ∂x ⎢ ∂ σ x ε ⎝ ⎠ c ⎠ ⎣ ⎦ ⎣ ⎝ ⎦ εC2i εεCi εCi C C C C ∂ ∂ ∂ ∂ i i i i ⎞ ρh ⎛ + − Cc 2 2 − Cc3 k ∂r ∂r ⎠ ci2 ci ⎝ ∂x ∂x (29) ⎜

⎟

Sw = −K G,H2O a w (pH2O − pH2O,G ) MH2O

(35)

2

where KG,H2O (kmol/kPa/m /s) represents the overall mass transfer coefficient of H2O. MH2O (kg/kmol) is the molecular weights of water·pH2O,G (kPa) is the partial pressures of H2O in the main gas phase·pH2O (kPa) represents the saturated vapor pressure of the water, which can be gained as Antoine equation:

The constants in the c 2 − εc model are taken as Cc0 = 0.11, Cc1 = 1.8, Cc2 = 2.2, Cc3 = 0.8, σc = 1.0 and σεc = 1.0 [27–29]. (4) NH3 mass fraction equation in the liquid phase

pH2O = 10(7.07406 − 1657.46/(T − 46.13)

(36)

The overall mass transfer coefficient of H2O can be represented by 5

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where MCO2 (kg/kmol) and MNH3 (kg/kmol) are the molecular weight of CO2 and NH3.

the gas phase mass transfer coefficient because of the ignoring of the liquid film resistance [30,31]. Therefore, KG,H2O can be expressed as follows [30–32]:

K G,H2O = k g,H2O =

2.4.4. Source term of rich solution The decreasing of the ammonium bicarbonate (NH4HCO3) in the rich solution is owing to the regeneration of ammonia. Therefore, the source term Sdes (kg/m3/s) can be expressed as follows:

μg 1 D a GM 0.7 ) (ρ D ) 3 (aT dp)−2 ( LR, wT T ) T μg g g g H2 O

5.23 ∗ ( a

(37)

3600 ∗ 101.325

where GM (kg/m /h) and μg(Pa·s) are the gas phase mass velocity and viscosity, respectively. The losing H2O of the liquid phase is the adding H2O of the gas phase. Therefore, the source term of H2O mass fraction in the gas phase can be obtained as follows: 2

SH2O = −Sw

Sdes = −

2.4.2. Source term of CO2 desorbing SCO2 (kg/m3/s) is the CO2 desorbed from the rich solution to the gas phase per unit volume and time. As the literature of Ma et al. [9], the desorption reaction is second-order reaction. Therefore, the desorbing rate of CO2 can be expressed as follows:

Sm = Sw − SNH3_G − SCO2

3

where XHCO−3 (kmol/m ) and XNH+4 (kmol/m ) are the molar concentration of HCNO3− and NH4+, respectively. MCO2 (kg/kmol) and kr (kmol/m3/s) represent CO2 molecular weight and reaction rate constant. kr (kmol/m3/s) can be obtained according to the Arrhenius equation [9,33].

kr = 4.4 × 1013 exp(

ST _G = hexc a w (T − TG ) + ∗ 106 +

)

(

ST _L = −

(41)

pNH3, i = HNH3 XNH3

hexc = kG (

where aw (m /m ) is the contact area of the gas and liquid phases in the stripper·pNH3 and pNH3,i (kPa) represent the NH3 partial pressures in the main gas phase and at the interface, respectively. HNH3 (kPa·m3/kmol) is the Henry’s constant of ammonia, which is used to predict the mass transfer of NH3 from liquid to gas phase in this paper. XNH3 and XNH3,i (kmol/m3) are the concentrations of ammonia in the main liquid phase and at the interface, respectively. kG,NH3 (kmol/m2/s/kPa) and kL,NH3 (m/s) are the gas phase and liquid phase mass transfer coefficients of NH3, which can be derived as the former literature of Chu et al. [28]. Owing to the high temperature in the stripper, the ammonia Henry’s constant is quite different to that in the absorbing columns. Therefore, it is important for the evaporation of NH3 to determine the Henry’s constant HNH3 (kPa·m3/kmol) correctly. The Henry’s constant of NH3 is expressed as follows [37,38]:

ρG (cp, G / MW , L ) λ2 DG2,H2O

1

)3

(49)

2.5. Boundary conditions The boundary conditions of the liquid and gas phase in the stripper are shown in Fig. 2, in which the mass and heat transfer process are presented in detail. The boundary conditions for the continuity, energy and all the mass conservation equations are presented as follows: (1) Inlet: The rich solution enters at the top of the stripper, and the boundary conditions are set as follows: U = Uinlet, Vinlet = 0, T = Tinlet, kinlet = 0.003(uinlet)2, εinlet = 0.09(kinlet)1.5/dH, Ci = Ci,inlet, CNH3 = C NH3,inlet, Cw = Cw,inlet.. The variance and dissipation rate equations are derived as Chu et al. [28,29]. (2) Outflow: The temperature of gas phase is set at this boundary, where the gas phase enters. Because the fully developed condition of the rich solution can be achieved at the bottom of stripper, the outflow boundary condition is consequently set

8621.06 −25.6767 ln T + 0.035388T ⎞ ∗ 103 T ⎠

SCO2 MNH3 MCO2

(48)

where λ (w/m/K) is the gas thermal conductivity. MW,L (kg/mol) is the average molar weight of the liquid phase.

(43)

As shown in Fig. 1, the generation of the ammonia in the liquid phase is due to the desorbing reaction of rich solution. At the same time, NH3 escapes from the liquid phase to gas phase. Therefore, the source term SNH3 (kg/m3/s) of ammonia in the liquid phase is calculated from the followings:

SNH3 = −SNH3_G +

|SCO2 | HR − ST _G MCO2

(42)

3

/ ρ ∗ 101.325

(47)

where HR (J/kmol) is the endothermic chemical reaction heat. HR = 6.487 × 107J/kmol. The heat transfer coefficient hexc (w/m2/K) is obtained by the Chilton–Colburn analogy, which is derived from Saimpert et al.[31]

)

SNH3_G = kG,NH3 a w pNH3, i − pNH3 MNH3 = kL,NH3 a w XNH3 − XNH3, i MNH3

HNH3 = exp ⎛160.559 − ⎝

SH2O ∗ H7 ∗ 106 MH2O

where hexc (w/m /K) is the convective heat transfer coefficient. The regeneration process of ammonia is an endothermic reaction, which is mainly supported by the rich solution. The energy source term of the liquid phase can be calculated as follows:

(40)

2.4.3. Source term of ammonia The NH3 mass fraction in the gas phase increases owing to the escape of NH3 from the liquid phase. Therefore, the source term can be calculated by the Two-Film theory [25,34–36]:

2

SNH3_G SCO2 ∗ H5 ∗ 106 + ∗ H6 MCO2 MNH3

2

−13541 ) T

(

(46)

2.4.5. Energy equation source term The heat exchange between liquid and gas phase consists of the physical volatilization heat and the convective heat transfer. Therefore, the source term gas phase energy can be simplified as follows:

(39)

3

(45)

where MNH4HCO3 (kg/kmol) represents the molecular weight of NH4HCO3. The loss of liquid mass is due to the volatilization of CO2, NH3 and H2O. The source term Sm (kg/m3/s) of the liquid continuity equation can be expressed as follows consequently:

(38)

SCO2 = MCO2·kr ·XHCO−3 XNH+4

SCO2 MNH4 HCO3 MCO2

∂Φ =0 ∂n

(50)

(3) Axis: The hydraulic, heat and mass transfer characteristics are assumed to axis-symmetrical, so the axis condition is set as follows:

(44) 6

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follows [12,39,40]:

q = H3 +

cp, r Δth φΔx

pH2O,top / pCO2 ,top

+ H7

(52)

where Δx = x1 − x2, x1 and x2 are the CO2 loadings of the rich and lean solution in the stripper, respectively. φ is the mass fraction of ammonia in the liquid phase before absorbing CO2·pH2O,top (kPa) and pCO2,top (kPa) are the partial pressures of water vapor and CO2 in the gas phase at the top of the stripper, respectively. The first part is the heat of the regeneration reaction (qreac); the second one is the sensible heat of the rich solution (qsens); the last one is the vaporization latent heat of liquid phase (qvap). The overall ammonia escape mass transfer coefficient of pack column is rarely researched. The overall mass transfer coefficient is deduced through the Two-Film Theory, which is shown in Fig. 3. The driving force of the ammonia escape is the concentration gradient between liquid and gas phase. The flux equation can be expressed as follows:

(

)

(

NA _NH3 = kL,NH3 a w XNH3 − XNH3, i = K G,NH3 pNH3, i − pNH3

)

(53)

where NA_NH3 (kmol/m /s) is the molar flux of NH3. 2

(

)

Ωd GyNH3 = ΩNA _NH3·a w ·dh

(54) 2

where dh is a differential height. G (kmol/m /s) and yNH3 (kmol/m2/s) are the gas phase molar flow rate and the volume fraction of NH3 in gas phase. Inserting Eq. (55) into Eq. (54) results in Eqs. (56)–(58) Fig. 2. Schematics of ammonia regeneration process in the stripper and corresponding boundary conditions.

∂Φ =0 ∂r

GdyNH3 = NA _NH3·a w ·dh

(55)

(

)

(56)

K G,NH3 a w ·p (yNH3, i − yNH3 )

(57)

GdyNH3 = K G,NH3 pNH3, i − pNH3 ·a w ·dh (51)

dh =

(4) Wall: The boundary at the wall is set the no-slip condition with the standard wall functions, which can predict the flow behavior near the wall in packing columns. (5) Model setting: all of the equations are solved by the Fluent. The SIMPLE and QUICK algorithms are selected. The realizable k-epsilon model is set. The convergence criterion for the maximum residual is set to 1 × 10−5.

GdyNH3

Top

∫Bottom dh =

G

Top

∫ K G,NH a w ·p Bottom (y

dyNH3

NH3, i

3

− yNH3 )

(58)

Eq. (59) is gained from the integral Eq. (58) from the bottom to the top of the stripper and the overall mass transfer coefficient can be obtained as follows [41,42]:

2.6. Performance parameters

K G,NH3 a w =

The specific energy consumption q (kJ/molCO2) of the ammonia regeneration process is composed by three parts, which is presented as

G H ·p

Top

∫Bottom (y

dyNH3

NH3, i

− yNH3 )

Fig. 3. Schematics of the ammonia escape mass transfer by the two-film theory. 7

(59)

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temperature in the stripper is presented in Fig. 5(b), which reveals that the temperature in the stripper is much higher than the absorber. The drive force of the ammonia escape is consequently larger than that in the absorber. As shown in Fig. 5(a), the escaped NH3 near the wall is higher than that at the axis, due to the higher temperature near the wall. Fig. 5(c) is the profile of the liquid temperature, which shows that the temperature in the center is a little higher than that near the wall. The reaction near the wall is not as fast as that in the center due to the wall flow effect. Therefore, the regeneration reaction in the center absorbs more heat than that near the wall.

2.7. Model verification The transport and hydromechanics performances are quite complex. Therefore, it is difficult to measure the species concentration and temperature distribution in the stripper. The numerical model in this paper is verified by two rate-based models [19,43], which were validated with the data of Munmorah power station. Yu et al. [43] reported the profiles of HCO−3 concentration in the liquid phase, which is used to be compared with the results in this paper. The height and diameter of the stripper are 3.5 m and 0.4 m, respectively. The ammonia concentration before absorption is 5% (wt). The stripping temperature of (a) 125 °C and (b) 115 °C. The flow rates of rich solution are (a) 60 L/ min and (b) 120 L/min, which are both with 0.4 mol/mol CO2 loading. The comparisons of the data from Yu et al. [43] and this numerical results are displayed in Fig. 4(a), which shows that the simulated results present a satisfactory agreement with the data from Yu et al. [43]. It can be seen that there is some differences in Fig. 4(a), which are mainly following reasons. First, the main absorbing product is assumed to be NH4HCO3 in this model without taking carbonate and carbamate into consideration. Therefore, the results in this paper is a little higher. Second, the rate-based model established by Yu et al. [43] is based on some assumptions in Aspen Plus, which can lead to some errors. Third, the mass transfer process inside the stripper cannot be presented by the Aspen Plus accurately, which is good at the whole process simulation. Moreover, the changing trends in Fig. 4(a) are in a satisfactory agreement, showing that the computational model is reliable enough to predict the multi-phase hydrodynamics and complex reaction of the regeneration process in the stripper on the other hand. The temperature profile in the stripper is reported by Zhang et al. [19], which is used to verify the reliability of the energy model in this paper. The height and diameter of the stripper packing are 4.5 m and 0.5 m. The inlet temperature and flow rate of rich solvent are 90 °C and 76 L/min. The ammonia concentration before absorption is 2.535 kmol/m3. The CO2 loading of the rich solution is 0.4 mol/mol. Fig. 4(b) shows the variations of the liquid phase temperature with height. It can be observed that although there are some differences between them in the middle of the stripper, the outlet temperatures are the same. The differences in the middle is mainly due to the assumption, which is that the endothermic regeneration reaction is only provided by the liquid phase. What’s more, the Aspen Plus is not good at the heat transfer process simulation inside the stripper. Therefore, the numerical model in this paper is reliable enough to predict the heat and multi-species transfer process in the stripper.

3.2. Rich solution flow rate and inlet temperature The rich solution flow rate and inlet temperature are two key operating parameters for the regeneration process. It is of benefit for the application to clarify the effects of rich solution flow rate and inlet temperature on the regeneration energy and ammonia escape behavior. The detailed information of the working conditions is listed in Table 1. The variations of the ammonia escape overall mass transfer coefficient with the inlet temperature under different rich solution flow rates are shown in Fig. 6, which illustrates that KG,NH3aw increases significantly with the rising of the liquid phase inlet temperature. The higher inlet temperature can result in a bigger Henry’s constant of ammonia, which can lead to the larger driving force of ammonia escape. The similar results were also reported by Wang [25] qualitatively. Moreover, KG,NH3aw varies little with the liquid flow rate change. It can be gained that the mass transfer performance of ammonia escape is mainly controlled by the liquid inlet temperature, rather than liquid flow rate. The variations of the specific regeneration energy consumption with the rich solution inlet temperature under different rich solution flow rates are shown in Fig. 7, which reveals that increasing the rich solution inlet temperature can lead to a significant reduction of the specific regeneration energy consumption. As the rich solution inlet temperature increase, temperature difference between the liquid and gas phases decreases. Therefore, the sensible heat of the rich solution decease and the higher average temperature can lead to the increasing of the desorbed CO2 amount. As shown in Fig. 7, the specific energy consumption decreases as the liquid flow rate increases slightly. The rising of the liquid flow rate can lead to the increase of the reactant, which is beneficial to the regeneration reaction. Ma et al. [10] reported the consistent researches with the results in this part. The variations of the three energy consumption parts in the regeneration with the rich solution flow rate at the 368.15 K inlet temperature are shown in Fig. 8(a) and the variations of the three energy consumption parts in the regeneration with the rich solution inlet temperature at 35 m3/m2/h liquid flow rate are shown in Fig. 8(b).

3. Results and discussion It is of significance to clarify the effects of working conditions on the ammonia escape performance and energy consumption for the ammonia technology application. An industrial scale stripper equipped with Sulzer Mellapak 350Y is investigated in this work, for which the height is 8 m and the diameter is 2 m. 3.1. Stripper inside view The species transfer from liquid phase to gas phase is quite complex and the regeneration reaction is too complicated. It is consequently impossible to get the profiles of all the species concentration and the temperature inside the strippers only through the experimental method. The CFD simulation has become the effective method to predict the ammonia regeneration process, which can provide the inside view of the stripper. The inlet flow rate and the temperature of rich solution are set to be 25 m3/m2/h and 353.15 K, respectively. The inlet rich loading and ammonia concentration are 0.6 mol /mol and 5 kmol/m3. The gas inlet flow rate and temperature are 2500 m3/m2/h and 368.15 K. The profile of the escaped NH3 in the gas phase is shown in Fig. 5(a), which is much high than that in the absorber [28]. The contour of gas

Fig. 4a. Comparisons between this simulation and the model by Yu. 8

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Fig. 4b. Comparisons between this simulation and the model by Zhang.

Fig. 5b. Contour of the gas temperature in the stripper.

Fig. 5a. Contour of the escaped NH3 in the gas phase.

Fig. 8(a) shows that the sensible heat of the rich solution (qsens) and vaporization latent heat of liquid phase (qvap) decrease a little as the rich solution flow rate rising. On the other hand, Fig. 8(b) illustrates that both qsens and qvap goes down rapidly with the rising of liquid inlet temperature.

Fig. 5c. Contour of the liquid temperature in the stripper. Table 1 Detailed information of the working conditions for Section 3.2.

3.3. Rich loading and ammonia concentration

Item

The rich loading and initial ammonia concentration before absorbing affect both the regeneration energy and ammonia escape behavior. The ammonia loss performance and regeneration energy consumption are investigated under different rich loading and ammonia concentration. The detailed information of the working conditions is listed in Table 2. Both the flow rates and inlet temperature of the liquid and gas phases are kept a constant, so the ammonia volume fraction in the gas phase at the outlet of the stripper can represent the ammonia escape performance consequently. The variations of the outlet NH3 volume fraction in gas phase with the rich loading under different ammonia concentration are shown in Fig. 9. The outlet ammonia volume fraction decreases with the rich loading rising. The higher CO2 loading means the low free ammonia concentration in the liquid phase, which can reduce the drive force of the ammonia escape. As shown in

Values

liquid flow rate (m /(m ·h)) rich solution inlet temperature (K) ammonia concentration before absorption (kmol/m3) CO2 loading (mol/mol) CO2 concentration in gas phase at the bottom (kmol/m3) gas phase inlet temperature (K) gas flow rate (m3/(m2·h)) 3

2

15–35 343.15–368.15 5.0 0.3 0 378.15 2500

Fig. 9, the descending tendency is more obvious at the higher ammonia concentration. The ammonia escape becomes more serious as the initial ammonia concentration increasing. The increasing of the initial ammonia concentration has the more obvious promoting effects than the rich loading. The variations of the specific regeneration energy consumption with 9

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Fig. 8b. Variations of q with rich solution inlet temperature at 35 m3/m2/h liquid flow rate.

Fig. 6. Variations of KG,NH3aw with inlet temperature under different liquid flow rate.

Table 2 Detailed information of the working conditions for Section 3.3. Item

Values

liquid flow rate (m /(m ·h)) rich solution inlet temperature (K) ammonia concentration before absorption (kmol/m3) CO2 loading (mol/mol) CO2 concentration in gas phase at the bottom (kmol/m3) gas phase inlet temperature (K) gas flow rate (m3/(m2·h)) 3

2

25 353.15 1.0–5.0 0.1–0.6 0 368.15 2500

Fig. 7. Variations of q with liquid inlet temperature under different liquid flow rate.

Fig. 9. Variations of yNH3,out with rich loading under different ammonia concentration.

unacceptable. The energy consumption decreases with the initial ammonia concentration and CO2 loading rising. Jilvero et al. [44] also reported that the heat requirement increase rapidly at a lower ammonia concentration. The tendency of CNH3 = 5 kmol/m3 is not obvious due to the range of left Y axis is too large. Therefore, it is presented in the blue line especially. The variations of the three energy consumption parts in the regeneration with the ammonia concentration at 0.6 rich loading are shown in Fig. 11(a) and the variations of the three energy consumption parts in the regeneration with the rich loading under 5kmol/ m3 ammonia concentration are shown in Fig. 11(b). As shown in Figs. 11(a) and 11(b), it can be seen that both the sensible heat and vaporization latent heat decrease with the increasing of the initial

Fig. 8a. Variations of q with rich solution flow rate at 368.15 K inlet temperature.

the rich loading under different initial ammonia concentration are presented in Fig. 10, which shows that the energy consumption at 1 kmol/m3 initial ammonia concentration is extremity high and

10

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Table 3 Detailed information of the working conditions for Section 3.4. Item

Values

liquid flow rate (m /(m ·h)) rich solution inlet temperature (K) ammonia concentration before absorption (kmol/m3) CO2 loading (mol/mol) CO2 concentration in gas phase at the bottom (kmol/m3) gas phase inlet temperature (K) gas flow rate (m3/(m2·h)) 3

2

25 353.15 5.0 0.3 0 358.15–378.15 1000–3000

Fig. 10. Variations of q with rich loading under different initial ammonia concentration.

Fig. 12. Variations of KG,NH3aw with gas flow rate under different inlet temperature.

Fig. 11a. Variations of q with initial ammonia concentration at 0.6 rich loading.

Fig. 13. Variations of q with gas flow rate under different inlet temperature.

gas flow rate and inlet temperature on the regeneration energy and ammonia escape behaviors. The detailed information of the working conditions is listed in Table 3. The variations of the ammonia escape overall mass transfer coefficient with the gas flow rate under different inlet temperature in Fig. 12, which illustrates that the gas inlet temperature has little impacts on the KG,NH3aw. On the other hand, the KG,NH3aw increases with the rising of the gas flow rate. The increasing of the gas flow rate can lead to a better contact of the liquid and gas phase, which is beneficial to the mass transfer of the ammonia escape. However, the range of Y axis is small, which shows that the effects of gas flow rate on the ammonia escape mass transfer are not obvious. Therefore, the mass transfer of ammonia escape is mainly controlled by the liquid film, rather than the gas film.

Fig. 11b. Variations of q with rich loading at 5kmol/m3 initial ammonia concentration.

ammonia concentration and CO2 loading. 3.4. Gas flow rate and inlet temperature The regeneration process is also affected by the condition of the gas phase. It is of significance for the application to study the impacts of the 11

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the liquid inlet temperature and initial ammonia concentration, which shows that the mass transfer of ammonia escape is mainly controlled by the liquid film. The increasing of the initial ammonia concentration and the liquid inlet temperature can accelerate the ammonia loss. Moreover, the regeneration energy consumption can be reduced by the increasing of the liquid inlet temperature, initial ammonia concentration and CO2 loading. Especially, the energy consumption of 343.15 K liquid inlet temperature is almost three times to that of 368.15 K liquid inlet temperature. The lower gas phase temperature and flow rate not only can inhibit the ammonia escape but also can reduce the regeneration energy consumption. Acknowledgments This is for you, my BUCT. The research is supported by the National Natural Science Foundation of China (Grant No. 51906012). The financial supports for this research, from the Fundamental Research Funds for the Central Universities (BUCTRC201922) and the National Natural Science Foundation of China (Grant No. 51776067), are also gratefully acknowledged.

Fig. 14a. Variations of q with the gas inlet temperature at 1000 m3/m2/h gas flow rate.

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Fig. 14b. Variations of q with the gas flow rate under gas inlet temperature 353.15 K.

The variations of the specific regeneration energy consumption with the gas flow rate under different gas inlet temperature are shown in Fig. 13, which illustrates the rising of the gas flow rate and inlet temperature can result in a remarkable increase of the energy consumption. The variations of the three energy consumption parts in the regeneration with the gas inlet temperature at 1000 m3/m2/h gas flow rate are shown in Fig. 14(a) and the variations of the three energy consumption parts in the regeneration with the gas flow rate under gas inlet temperature 353.15 K are shown in Fig. 14(b). The rising of the gas inlet temperature can lead to the increase of the temperature difference between liquid and gas phase, which can liquid to the increase of the sensible heat of the rich solution as shown in Fig. 14(a). Fig. 14(b) reveals that the vaporization latent heat increases obviously with the rising of the gas flow rate. The increasing of the gas flow rate leads to more heat taken out of the stripper. Therefore, the lower gas phase condition is of benefit for both the ammonia escape control and regeneration energy saving. 4. Conclusions The mass transfer performance of ammonia escape and the accurate energy consumption of regeneration are both achieved in this paper based on a novel ammonia regeneration model. The ammonia escape performance in the ammonia regeneration process is greatly affected by 12

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