Numerical Analysis of the Flow Characteristics and Heat and Mass

Article pubs.acs.org/IECR

Numerical Analysis of the Flow Characteristics and Heat and Mass Transfer of Falling-Water Films in an Industrial-Scale Dip Tube of a WSCC in an OMB Gasifier Yifei Wang,* Qiangqiang Guo, Bihua Fu, Jiangliang Xu, Guangsuo Yu, and Fuchen Wang Key Laboratory of Coal Gasification and Energy Chemical Engineering of the Ministry of Education, East China University of Science and Technology, Shanghai 200237, PR China ABSTRACT: The flow, heat, and mass transfer of falling-water films in a dip tube in an industrial-scale water-scrubbing cooling chamber was predicted using a 2D numerical simulation. The VOF model was adopted to describe the free surface between the gas and liquid. The Rosseland model was applied to solve the radiative heat-transfer equation. The surface tension between the gas and liquid phases was also considered. The predicted axial-temperature distribution of the bench-scale results were in good agreement with the experimental data, validating the feasibility of the numerical models applied in this Article. The simulation results indicated that the thickness of the inlet water film and the velocity of the inlet falling-water film in the industrial-scale dip tube had a large influence on the mass, heat transfer, and flow patterns of the falling-water film. When the thickness of the inlet water film was increased to 0.004 m, the outlet hot-syngas temperature of the dip tube decreased to about 100 K. Furthermore, the results showed that the hot syngas was humidified sufficiently during the mass-transfer process.

1. INTRODUCTION A water-scrubbing cooling chamber (WSCC) is applied to an opposed multi-burner (OMB) gasifier1−3 in which hightemperature syngas (with molten slag) is cooled, washed, and humidified, and the coarse particles are collected. Although the recovery of sensible heat from gasification products by water quenching is less efficient than that from a water heating boiler, the direct water-cooling process is able to handle sticky particulates well and with a lower cost; more importantly, there could be enough steam used for the manufacture of methanol or oxo products, producing hydrogen for refineries and ammonia synthesis.4−6 As one of the most important items in a WSCC, the dip tube performs the tasks of hot-gas cooling and humidification. The heat- and mass-transfer characteristics of a falling-water film flowing along the dip tube wall are very important because of their influence on the outlet temperature of the dip tube, which eventually impacts the performance of the WSCC system. However, according to a stability theory for waves, vertical-falling film flows are always unstable.7 There are almost no reports on the influence of a vertical falling film on the temperature distribution within the hot syngas along the dip tube. The purpose of this Article is to develop a numerical model of multiphase flow with complex direct-contact heat and mass transfer in the dip tube where a large temperature difference is difficult to measure by experimental methods without disturbing the flow field. It is well known that the heat and mass transfer of a liquid falling film is enhanced by pressure fluctuations on the surface that follow the evaporation. Many experimental and numerical studies have been conducted to investigate the heat- and masstransfer characteristics of a falling film-flow field.8−10 However, most of these were focused on laminar-liquid falling film flows, whereas the mechanism of the heat-transfer enhancement resulting from a liquid falling film is still unclear. © 2013 American Chemical Society

The numerical simulation of interphase mass transfer can be performed by two fundamentally different approaches.11 Many reports12−14 revealed that the VOF method is one of the most widely used interface-tracking methods, whereas the others include level set, front tracking, and so on. The advantages15 of this method are that no redistribution of the surface markers is necessary when they are stretched by the flow and no special provision is necessary to perform a reconnection of the interfaces; however, this may be a disadvantage if one wishes to prevent reconnection from occurring. The VOF model16 was validated to simulate film boiling in multiphase-flow systems such as gas−liquid flow. As for the heat- and mass-transfer characteristics in the dip tube of a WSCC, the research has usually relied on a number of numerical simulations. Examples of such simulations can be found in the work of Wu et al.17 who simulated heat and mass transfer at a high pressure in WSCC by using the Runge−Kutta method. The temperature distribution of the hot syngas in the dip tube of a Texaco gasifier under different conditions, including pipe length, pipe diameter, inlet gas velocity, and inlet gas temperature, was predicted by using the VOF model and the finite-volume method (FVM).18,19 Li et al.20 performed a numerical simulation by using the VOF method and reported that the Rosseland model is more reliable than the other three radiative models (the discrete transfer radiation model, discrete ordinates, and P1) on an industry-scale dip tube in a Texaco gasifier. The simulation21 of heat and mass transfer on a benchscale dip tube in an OMB gasifier showed that the temperature of the syngas declined sharply in the upper half of the dip tube, allowing the optimum volume of cooling water to be predicted. Received: Revised: Accepted: Published: 9295

November 4, 2011 December 24, 2012 June 12, 2013 June 12, 2013 dx.doi.org/10.1021/ie2025294 | Ind. Eng. Chem. Res. 2013, 52, 9295−9300

Industrial & Engineering Chemistry Research

Article

∂ (ρE) + ∇(υ ⃗(ρE + p)) = ∇(keff ∇T ) + S h ∂t

In the present study, a 2D numerical simulation of heat and mass transfer in an industrial-scale dip tube of a WSCC in an OMB gasifier was performed to investigate how they were influenced by the thickness and velocity of an inlet falling-water film and water-vapor distribution in the dip tube. The hot syngas-temperature distribution in dip tube of bench-scale WSCC was measured and the comparison of the predicted results and measured data were also carried out.

where the VOF model treats enthalpy, E, as a mass-averaged variable:

Tsat

−γρl (T − Tsat) ∂ml = ∂α Tsat

γρl α(T − Tsat) Tsat

∂ṁ v =0 ∂α

(11)

∂ ∂ (ρκ ) + (ρκui) ∂t ∂xi ∂uj ⎞ μ ⎞ ∂κ ⎤ ∂u ⎛ ∂u ∂ ⎡⎢⎛ ⎟ − ρε = ⎜μ + t ⎟ ⎥ + μt i ⎜⎜ i + ∂xi ⎟⎠ σκ ⎠ ∂xj ⎥⎦ ∂xj ⎢⎣⎝ ∂xj ⎝ ∂xj

⎛ ∂υ ⎞ + υ∇υ⎟ = −∇P + ∇[μ(∇υ + ∇υT )] + ρg ρ⎜ ⎝ ∂t ⎠ (3)

Because of smoothing, the surface-tension force is applied to a transition region a few cells thick centered at the interface. The curvature calculation is implemented using second-order central differences, and a discussion of the effects of smoothing and accuracy can be found in ref 23. The density and viscosity vary with the volume fraction as

(5)

(10)

2.4. Turbulent Model. The standard κ−ε model is the simplest turbulence model proposed by Launder and Spalding for practical engineering flow calculations.25 The turbulence kinetic energy κ and its rate of dissipation ε are shown as follows

(2.2.2). Momentum Equation. The momentum equations are augmented using the continuum-surface-force model of Brackbill et al.23

μ = αlμ l + (1 − αl)μg

(9)

The mass transfer from water, ṁ l, results from the vapor enhancing the mass in the ullage volume. The vapor mass transport is governed by the following relations:

(1)

(4)

(8)

The source term ṁ l is taken to depend only on one solution variable, the water volume fraction α. γ is the time-relaxation factor used for obtaining the rate of mass transfer in a cell of unit volume. The derivative with respect to α is therefore

ṁ v =

ρ = αlρl + (1 − αl)ρg

(7)

−γρl α(T − Tsat)

ṁ l =

(2)

+ σκc∇α̃

αlρl + αgρg

In the previous investigation, the Rosseland model20 was recommended to use for simulating radiation heat transfer in a vertical pipe. 2.3. Mass Transfer. The interfacial mass transfer ṁ represents the mass lost by the liquid phase due to evaporation at the interface, with the converse being true for the vapor phase. The interfacial mass and heat transfer resulting from evaporation are modeled in a manner inspired by Kumar et al.24 Evaporation takes place when the enthalpy of the liquid exceeds the saturated value. The water-mass transfer is given by

The volume fraction equation will not be solved for the syngas phase (primary phase), and the syngas-phase volume fraction is computed on the basis of the following constraint: αl + αg = 1

αlρl E l + αgρg Eg

E=

2. MODEL DESCRIPTION 2.1. Fluid Flow Approach. The multiphase flow in the dip tube is unsteady and does not interpenetrate. The following assumptions are made to simplify the description of the gas− liquid two-phase flow in the dip tube. (1) The syngas is incompressible. (2) The inlet velocities of the syngas and water are constant. (3) The influence of ash on the heat transfer is ignored. (4) The temperature of the wall of the dip tube, which is the saturated temperature of the water under the corresponding vapor pressure, is steady. The volume of fluid (VOF) method, validated by Welch and Wilson,22 is a viable option for the simulation of gas−liquid flows, including cases with mass transfer. Hence, it is used to describe the free surface between the water and syngas in this Article. To track the moving gas−liquid interface, the volume fractions of the liquid and syngas, αl and αg, are introduced. In each control volume, the sum of the volume fractions of the two phases is unity. On the basis of the local value of αl, the appropriate properties and variables will be assigned to each control volume within the domain. 2.2. Governing Equations. (2.2.1). Continuity Equation. The tracking of the interface between the syngas and water is accomplished by the solution of a continuity equation for the volume fraction of the second phase (the liquid phase is defined as the second phase). ⎡ ∂α ⎤ ρl ⎢ l = ∇(υ ⃗ ·αl)⎥ = ṁ ⎣ ∂t ⎦

(6)

(12)

and ∂ ∂ (ρε) + (ρεuj) ∂t ∂xj =

∂uj ⎞ μ ⎞ ∂ε ⎤ C ε ∂ui ⎛ ∂ui ∂ ⎡⎢⎛ ⎜⎜ ⎟ + ⎜μ + t ⎟ ⎥ + μt 1 ∂xi ⎟⎠ ∂xj ⎢⎣⎝ σε ⎠ ∂xj ⎥⎦ κ ∂xj ⎝ ∂xj − C2ρ

(2.2.3). Energy Conservation Equation. 9296

ε3 κ

(13)

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where the standard constants of the standard κ−ε model used were given as Cμ = 0.09, C1 = 1.44, C2 = 1.92, σκ = 1.0, σε = 1.2.

FVM. The velocity inlet and pressure outlet were used for the inlet and outlet boundaries, respectively. The velocity correction was realized to satisfy continuity using the SIMPLE algorithm,26 which couples the velocity and pressure. The body-force-weighted scheme was used for pressure discretization. The volume fraction was solved using Geo-Reconstruct, which describes the piecewise linear-interface reconstruction in Fluent software, and the second-order upwind scheme was used for the other terms (momentum, turbulence kinetic energy, turbulence dissipation rate, and energy). The convergence criteria for the flow and heat transfer (energy, velocity, continuity, κ, and ε) were achieved when the iteration residuals were reduced to 5 orders of magnitude. A constant temperature, which was the saturation temperature of water at the operational pressure, was applied to the wall. The numerical simulations were performed for an industrialscale dip tube, and the operating conditions of the calculation cases for the industrial-scale model are given in Table 1. The thickness, δ, (Figure 2) of the falling-water film at the inlet of the dip tube is a key factor that influences the terminal-gas temperature at the outlet of the dip tube. Therefore, the effects of the thickness, δ, for three cases, δ = 0.006 m (a), 0.008 m (b), and 0.010 m (c), on the falling water flow field and heat transfer were discussed.

3. BOUNDARY CONDITIONS AND SIMULATION METHOD The configuration of the dip tube is displayed in Figure 1. The 2D axisymmetric model of the domain (Figure 2) is meshed

Figure 1. Configuration of the dip tube.

4. HOT-MODEL EXPERIMENT AND MODEL VALIDATION It is necessary to verify the accuracy of the computational scheme before it is applied to the industrial-scale-model prediction. The predicted temperature-distribution results were compared with the experimental data measured in the bench-scale hot-model system. The details of the OMB gasifier hot-model experiment have been reported by Niu et al.27 and Yan et al.28 Figure 3 is the schematic diagram of the hot-modelexperimental system. The cooling water supplied from the pump is injected into a cooling ring, flowed out from a 3 mm slit under the cooling-ring, and goes down as a falling-water film along the wall of the dip tube. The schematic diagram of the temperature test system is shown in Figure 4. The diameter and length of the bench-scale dip tube that is fixed in the WSCC and vertically oriented are 0.1 and 1 m, respectively. There are four thermocouples fixed on the central axis of the dip tube, and to avoid the influence of water or ash, the thermocouples are covered by a metal sleeve. The axial temperature the of dip tube is obtained by adjusting the location of the four thermocouples. The signals from the thermocouples are transferred into a computer by ADAM which then uses the MCGS configuration module to test the axial temperature. The inlet gas temperature is 1100 k, the inlet gas velocity is 1.86 m·s−1, and the quenching-water temperature and velocity are 280 K and 0.55 m·s−1, respectively. The comparison of the experimental data and the predicted axial-temperature distribution of the dip tube is shown in Figure 5. The results showed that the hot syngas was rapidly quenched and that the syngas temperature had decreased by more than 740 K at the starting point (∼0.2 m) of the bench-scale dip tube. Meanwhile, a fierce heat and mass transfer occurred

Figure 2. Schematic diagram of the 2D physical models.

with the structured grids of the quad mesh. Furthermore, the properties of the syngas and water vary by several orders of magnitude. The local mesh is refined using a series of successively increasing meshes located toward the wall to capture the steep flow gradients of the falling-water film and the syngas temperature at the interface. We provide 500 × 30 600 × 40 800 × 400 three-type grids. The specified grid is fine enough to give a grid-dependent solution and can be validated through the grid-dependent tests. The grid-sensitivity tests indicated that a hybrid mesh with 24 000 elements is sufficient to obtain a grid-independent solution. The governing equations for the conservation of momentum, energy, mass, and radiation were solved sequentially by the Table 1. Simulation Conditions of the Industrial-Scale Dip Tube D

H

quenching water velocity

quenching water temperature

inlet gas temperature

inlet gas velocity

pressure

0.86 m

4m

1−4 m·s−1

500 K

1700 K

5 m·s−1

3.0 MPa

9297

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multiphase flow or heat- and mass-transfer processes of the industrial-scale dip tube of the WSCC can be applied to an OMB gasifier.

5. RESULTS AND DISCUSSION 5.1. Temperature Distribution. The thickness of the inlet water film and the velocity of the inlet cooling water are two of the main factors influencing the hot syngas-temperature distribution of the gas−water two-phase flow in the industrial-scale dip tube of the WSCC applied in an OMB gasifier. The total cost of the WSCC system is also an important factor. The temperature-distribution profiles of the hot syngas under three calculation conditions were compared and shown in Figure 6. The three conditions of the thicknesses are a, b, and c (detailed in section 3), with each having constant cooling water and syngas velocities. The obvious temperaturedecreasing processes were taking place in the first half of the dip tube (∼1 to 2 m) and were not affected by the water-film thickness. At the same time, strenuous heat and mass transfer had occurred, and most of the cooling water had been evaporated into the vapor phase, which will be discussed later. Additionally, in practical industrial operation the first half of the dip tube is damaged in many cases, which proved that the numerical models discussed above were reasonable. In Figure 6 (2, right panel), for hot syngas the calculated syngas temperature of the dip tube outlet was significantly affected by the thickness of the inlet water film. When δ = 0.010 m, the hot syngas temperature decreased from more than 1700 K to less than 750 K. However, the outlet temperature of the dip tube was about 850 K when δ = 0.0060 m, and as the fallingfilm thickness decreased to 0.004 m, the outlet temperature of the dip tube increased by about 100 K. The hot-syngas-temperature distribution along the dip tube at different inlet water velocities is shown in Figure 7. In this case, the gas velocity of 3 m·s−1 and the inlet water velocities u1, u2, u3, and u4 of 1, 2, 3, and 4 m·s−1, respectively, were kept constant. It was also found that the main cooling process ended at about 2 m, as displayed in Figure 6. The outlet temperature of the dip tube decreased rapidly with the increasing inlet water velocities, which means that the total volume of water increased. When the water velocity was 1 m·s−1, the outletsyngas temperature of the dip tube was very high, ∼1000 K, and the temperature increased slightly from 2 m to the end of the dip tube, as displayed in Figure 7b, which demonstrated that most of the water had been evaporated at the end of the 2 m and the total volume of water was not enough to cool the syngas. However, the outlet temperature of the syngas dropped from 1700 K to about 700 K, as displayed in curve u4. The outlet-temperature difference of curves u3 and u4, which was ∼50 K, is half of that of u2 and u3 (i.e., ∼100 K). Although the total volume of the cooling water was influenced significantly by the outlet temperature of the hot syngas of the dip tube, an optimum condition still existed. As demonstrated, the increased water velocity could not promptly decrease the outlet-syngas temperature when the velocity exceeds the maximum limit, and most importantly, merely increasing the water volume is not economical or viable because it would increase the loads of the subsequent processes, such as the entrained water. Effectively, at the industrial-scale the dip tube length is about 4 m, and the measured syngas temperature at the outlet of the dip tube is about 600 °C.29 Hence, it can be deduced that the optimum condition is when the cooling-water film thickness is δ = 0.006 m or the cooling-water velocity is 2 m·s−1.

Figure 3. Schematic diagram of the hot-model-experimental system. The components of the system are (1) the scrubbing-cooling chamber, (2) the cooling-water inlet, (3) the gasifier segment, (4) the water inlet, (5) the water jacket, (6) the nozzle, (7) the flame-vision lens, (8) the water outlet, (9) the nozzle-position sketch map, (10) the gas outlet, (11) the cooling-water inlet-position sketch map, (12) the cooling-water outlet, and (13) the slag outlet.

Figure 4. Schematic diagram of the temperature-test system The components of the system are (1) the metal sleeve of the thermocouple, (2) the thermocouple, (3) the dip tube, (4) the scrubbing-cooling chamber, (5) ADAM, and (6) the computer.

Figure 5. Comparison between the measured and predicted axialtemperature profiles.

during the period of the temperature decrease. In Figure 5, it can also be seen that the predicted results agree well with the experimental data. Thus, it was proven that the numerical models discussed above were reasonable. Hence, on the basis of the above models, the following numerical prediction for the 9298

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Figure 6. Comparison of the axial-temperature distribution of the dip tube at different inlet water film thicknesses at an inlet water velocity of 3 m·s−1.

Figure 7. Schematic diagram of the axial-temperature distribution of the dip tube at different inlet water velocities at an inlet water-film thickness of 0.006 m.

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6. CONCLUSIONS

AUTHOR INFORMATION

Corresponding Author

A 2D-CFD model has been developed to investigate the heatand mass-transfer processes and characteristics of a falling-water film under different simulation conditions in an industrial-scale dip tube in a WSCC of an OMB gasifier. The axial-temperature distribution and water film patterns were obtained by numerical simulation and experimental research. The major conclusions are summarized as follows: (1) The comparison between the calculated results and the experimental data revealed that the numerical models discussed in this Article were appropriate for studying heat- and mass-transfer characteristics as well as the falling film-flow behaviors in an industrial-scale dip tube. (2) The inlet water-film thickness had a great effect on the axialtemperature distribution and flow patterns of the falling-water film. A remarkable temperature drop occurred after 2 m, with an approximately 100 K decrease in the final temperature at the dip-tube outlet when the difference between the thickness of the two films was 0.004 m. (3) The inlet water-film velocity had a significant effect on the heat and mass transfer and the fallingwater-film flow characteristics in the dip tube. When the velocity was less than 2 m·s−1 there was not enough cooling water to cool the hot syngas, the outlet temperature was very high, and the water-film flow pattern was discontinuous.

*Tel: +86 21 6425 2522. Fax: +86 21 6425 1312. E-mail: [email protected]. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS This work is financially supported by the National Key State Basic Research Development Program of China (973 program, 2010CB227 000) and the National High-Tech Research and Development Program of China (2007BAA08B01).

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9299

NOMENCLATURE C1, C2 = constants in production, sink terms of e equation Cμ = constant in the constitutive equation of the κ−ε model D = diameter of dip tube (m) E = total energy (J) Eg = energy of the gas phase (J) El = energy of the water phase (J) H = height of the dip tube (m) m = mass transfer at the surface (kg) Sh = source term Sαl = user-defined mass source for the water phase T = temperature of gas (K) ug = velocity of gas (m·s−1) dx.doi.org/10.1021/ie2025294 | Ind. Eng. Chem. Res. 2013, 52, 9295−9300

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ui = velocity of water (m·s−1), i = 1, 2, 3, and 4

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Greek Symbols

α̃ = smoothed void fraction αg = volume fraction of the gas αl = volume fraction of the liquid μ = total viscosity (kg·m−1·s−1) μl = liquid viscosity (kg·m−1·s−1) μg = gas viscosity (kg·m−1·s−1) ρ = total density (kg·m−3) ρg = density of the gas (kg·m−3) ρl = density of the water (kg·m−3) γ = water evaporation factor κc = curvature of the surface σκ, σε = turbulent Prandtl number for the diffusion of κ and ε

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