Removal from Flue Gas with Ca(OH) - American Chemical Society

Subscriber access provided by - Access paid by the | UCSB Libraries

Article

SO3 removal from flue gas with Ca(OH)2 in entrained flow reactor Hui Wang, Denggao Chen, Zhenshan Li, Dinghai Zhang, Ningsheng Cai, Jin Yang, and Geng Wei Energy Fuels, Just Accepted Manuscript • DOI: 10.1021/acs.energyfuels.8b00274 • Publication Date (Web): 27 Feb 2018 Downloaded from http://pubs.acs.org on March 8, 2018

Just Accepted “Just Accepted” manuscripts have been peer-reviewed and accepted for publication. They are posted online prior to technical editing, formatting for publication and author proofing. The American Chemical Society provides “Just Accepted” as a service to the research community to expedite the dissemination of scientific material as soon as possible after acceptance. “Just Accepted” manuscripts appear in full in PDF format accompanied by an HTML abstract. “Just Accepted” manuscripts have been fully peer reviewed, but should not be considered the official version of record. They are citable by the Digital Object Identifier (DOI®). “Just Accepted” is an optional service offered to authors. Therefore, the “Just Accepted” Web site may not include all articles that will be published in the journal. After a manuscript is technically edited and formatted, it will be removed from the “Just Accepted” Web site and published as an ASAP article. Note that technical editing may introduce minor changes to the manuscript text and/or graphics which could affect content, and all legal disclaimers and ethical guidelines that apply to the journal pertain. ACS cannot be held responsible for errors or consequences arising from the use of information contained in these “Just Accepted” manuscripts.

Energy & Fuels is published by the American Chemical Society. 1155 Sixteenth Street N.W., Washington, DC 20036 Published by American Chemical Society. Copyright © American Chemical Society. However, no copyright claim is made to original U.S. Government works, or works produced by employees of any Commonwealth realm Crown government in the course of their duties.

Page 1 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

SO3 removal from flue gas with Ca(OH)2 in entrained flow reactor Hui Wang1, Denggao Chen1, Zhenshan Li1*, Dinghai Zhang2, 3, Ningsheng Cai1, Jin Yang2, 3, Geng Wei2, 3 1 Key Laboratory for Thermal Science and Power Engineering of Ministry of Education, Department of Energy and Power Engineering, Tsinghua University, Beijing 100084, China 2 Key Laboratory for Clean Combustion and Flue Gas Purification of Sichuan Province, Chengdu 611731, China 3 Dongfang Electric Group Dongfang Boiler Co. Ltd., Zigong 643001, China

Abstract: Experiments were carried out in a pilot-scale entrained flow reactor (EFR) to investigate the reaction of SO3 with Ca(OH)2, as a method of dry sorbent injection (DSI) for SO3 removal from flue gas. The results indicate that SO3 can be removed by Ca(OH)2, with an efficiency that can reach 80%, and it was found that the molar ratio of Ca(OH)2 to SO3 ([Ca]/[S]) and reaction temperature have a significant effect on SO3 removal efficiency. The experimental data measured inside the EFR were analyzed with a computing fluid dynamic (CFD) simulation in which the Euler-Lagrangian frames were used for gas and discrete phase modeling. The CFD models were validated and applied to analyze the effects of certain parameters on SO3 removal efficiency, such as particle velocity, [Ca]/[S], and temperature and residence time. It was found that the sorbent diameter has a significant influence on SO3 removal efficiency, with an obvious

ACS Paragon Plus Environment

1

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 2 of 30

decrease in efficiency if the Ca(OH)2 particle diameter increases. For example, if the sorbent diameter increases from 3 to 10 μm, the SO3 removal efficiency at the reactor outlet will decrease from 99% to 55%. A detailed comparison and theoretical analysis indicated that external diffusion of SO3 from the gas phase to the particle surface is the rate controlling step for larger Ca(OH)2 particles, and more attention should be paid to the competition between external diffusion and surface reaction when applying the DSI method for removing SO3 from flue gas.

Key words: SO3 removal, dry sorbent injection, CFD simulation, controlling mechanism

1. Introduction In coal-fired power plants, SO3 in flue gas arises from both combustion processes and the selective catalytic reduction (SCR) system.

1–2

Most of the sulfur in the fuel is oxidized to SO2,

and a small percentage, usually 0.1 to 1.0%,

3

can be further oxidized to SO3. There is a

significant increase in SO2/SO3 conversion through catalytic oxidation of SO2 in SCR systems, 4– 8

which almost doubles the SO3 content in the final flue gas. Higher levels of SO3 in the flue gas

of coal-fired boilers adversely affects numerous aspects of plant operation and performance,

9–14

including severe corrosion of back-end equipment, fouling of the air preheater (APH), and blocking problems in the downstream equipment. From an environmental point of view, SO3 forms fine sulfuric acid mist in flue gas desulfurization (FGD) systems, and the formed acid aerosol sizes are particularly small, with mean diameters below 3 μm,

15

which cannot be

efficiently removed in desulfurization scrubbers. 16 The acid aerosols will subsequently become a visible, trailing blue plume from the stack,

17

which may result in a severe impact on the

environment and human health. Therefore, it is essential to take measures to mitigate elevated SO3 amounts in coal-fired power plants.


2

Page 3 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

Sorbent injection technologies using alkaline-based reagents offer a promising method for SO3 mitigation. 18–21 Depending on the reagent used, they may be injected in a dry or wet form. Wet sorbent injection (WSI) is more complex than dry sorbent injection (DSI), and the operation, maintenance and capital costs are higher, 22 while the DSI method is relatively simple and easily implemented. The injection location is typically ahead of the APH, where the sorbent can reduce the total SO3 formed in both the boiler and SCR, thereby reducing the existing SO3 inside the APH, and it could minimize the fouling impact of ammonium bisulfate on APH and corrosion on the downstream equipment. In the DSI process, gas-solid mixing and reaction are the two critical steps for SO3 removal from the flue gas. It has been found that appropriate sorbent distribution and mixing is important for achieving high removal efficiency,

20–21

and effective mixing

between flue gas and sorbent particles can be achieved by designing the feeding system of sorbent particles optimally. Gas-solid reaction is very important for achieving high SO3 removal efficiency and maintaining a low cost. It has been demonstrated that numerous alkaline-based reagents can be utilized as a sorbent, including magnesium oxide, ammonia,

26–27

sodium bicarbonate,

28

and trona.

22

23

calcium hydroxide,

24–25

However, SO3 removal efficiency and cost

vary among different sorbents. The SO3-sorbent reaction has been studied and reported in the literature, and related research has focused on the following aspects: (i) sorbent performance comparison among different alkaline-based reagents,

23, 25, 29–30

including reaction behavior, cost,

and effects on the plant and environment; (ii) investigation of the effect of operational conditions on SO3 removal efficiency,

20–21, 28, 31

including temperature, sorbent and SO3 molar ratio, and

flue gas residence times. It was reported that hydrated lime (Ca(OH)2) has been successfully utilized as a sorbent for mitigating SO3, due to the relatively low cost and good reactivity. 24–25 Sorbent reactivity is commonly believed to have a significant effect on the SO3 removal


3

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 4 of 30

efficiency; therefore, a great deal of effort has been made to improve sorbent performance. However, it should be noted that the SO3 concentration in flue gas is very low and within the range of 10 to 60 ppm,

13

and the SO3 removal process is controlled not only by the gas-solid

chemical reaction, but also the external gas diffusion. Thus, understanding the controlling steps of SO3 removal is significant in improving the SO3 removal efficiency. The positive influence of injecting sorbent in SO3 removal has been extensively supported; however, the effect of the controlling mechanism of sorbent on SO3 removal has not been clarified. Accordingly, this problem is investigated in our study. This study aims to: (1) design and construct an entrained flow reactor (EFR) to investigate SO3 removal efficiency using a dry sorbent injection of Ca(OH)2 particles under different temperatures and [Ca]/[S] levels; (2) analyze the effect of operational parameters and sorbent properties on SO3 removal efficiency with CFD models; and (3) discuss the competition between the gas-solid reaction and external diffusion, and identify the rate controlling step. The results may aid in further developing high sorbent reactivity improving DSI performance.

2. Experiment An SO3 removal experiment was conducted using the experimental system illustrated in Figure 1. The experimental system consisted of an air supply unit, air preheater system, SO3 generating system, sorbent injection system, EFR, and sampling system. The experimental process was conducted as follows. (1) Compressed air was carefully measured and controlled at the designed flow rate in the air supply system. (2) Part of the supplied air was heated by the preheater system, to the SO3 generator temperature. (3) Hot air from the preheater was mixed with SO2 gas from the SO2 tank, and the SO2 flow rate was controlled by a mass flow controller in order to ensure a constant SO2 concentration in the mixed air. (4) The SO2-air gas was induced


4

Page 5 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

into a vanadium-based catalyst containing V2O5 in order to generate a gas mixture containing SO3. (5) Cold air from the air supply system was induced into the hot gas mixture to adjust the gas temperature, and finally, the mixed gases were injected into the EFR. (6) Using the dry sorbent injection method, Ca(OH)2 particles were injected, dispersed, and mixed with the gas mixture at the reactor inlet, and the injector was specially designed to ensure stability of the injected sorbent stream. (7) When the gas mixture containing SO3 and sorbent moved down the EFR, the SO3 was adsorbed and its concentration decreased. The nominal diameter of the sorbent injector tube was 20 mm, the entire length of the EFR was 6000 mm, and the nominal diameter was 200 mm. Several operational parameters used in the experiment are displayed in Table 1.

Figure 1. Schematic of SO3 removal experiment system.

Table 1. Operational parameters used in experiment


5

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 6 of 30

Operational parameter

Value

Air flow rate [Nm3/h]

150

Carrier gas flow rate [Nm3/h]

6.6

Sorbent particle size [μm]

~6

[Ca]/[S]

5~10

Temperature of reactor [oC]

290~400

SO3 concentration at reactor inlet [ppm]

~16.15

The SO3 concentration was measured at one-meter intervals along the reaction distance, so that values could be obtained during different reaction stages. Figure 2 provides a schematic diagram of the sampling system. A controlled condensation method (GB/T 21508-2008) was used to measure the SO3 concentration in the sample. Samples were collected through the sampling tube and heated steam was used to ensure that the sampling tube gas temperature was maintained above the acid dew point temperature. The particulate matter in the sample could be removed by the heated quartz filter. Thereafter, the sample passed through a spiral glass tube and the SO3 could be condensed into a sulfuric acid mist on glass wool, which was stuffed into the glass tube. The spiral glass tube was placed in the water bath during the sampling process. After passing the glass tube, the sample moved through the washing and separating bottles. Deionized water was used to convert the sulfuric acid mist in the spiral glass tube to SO42- in the flushing fluid, following which the spectrophotometer was used to determine the SO42- concentration in the flushing fluid. The SO3 concentration in the flue gas could be calculated from the SO42concentration in the flushing fluid and the sample gas volume.


6

Page 7 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

1-gas sampling tube; 2-heated filter; 3-spiral glass tube; 4-water bath; 5-washing bottle; 6-separating bottle; 7-wet flow meter; 8-pump; 9-baffle plate Figure 2. Schematic of sampling system.

The SO3 generation system stability test was carried out prior to the SO3 removal test. The SO3 concentration in the flue gas without sorbent feeding, obtained during seven tests, is illustrated in Figure 3. It can be observed that the SO3 generation results were almost stable. That is to say, the SO3 generation method using a vanadium-based catalyst was reliable. In the following SO3 removal test, the SO3 concentration at the reactor inlet was measured three times and the average value was set as the inlet SO3 concentration.

Figure 3. Stability test results of SO3 generation system (temperature of catalyst section is 400 oC).


7

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 8 of 30

3. Model The CFD method is a useful tool for simulating the SO3 removal process inside the EFR. Compared to experiments, CFD simulation can investigate the effects of operational parameters and sorbent properties on SO3 removal efficiency more thoroughly, which may provide further insights for optimizing SO3 removal experiments. In this study, the experimental data measured inside the EFR were analyzed using CFD simulation, in which the Euler-Lagrangian frames were used for gas and discrete phase modeling. 3.1 Gas phase model The gas phase model is developed based on the conservation equations of mass, species, momentum, and energy, as well as the equation of the state of ideal gas, as follows:

        Sm t

(1)

 YSO3     YSO3    J SO3  SSO3 t

(2)

           p       Smv t

(3)

     e        e  p       kT   h j J j      Srad  Sh t j  

(4)

p   RT ,

(5)









where  , YSO3 , J SO3 , p, T , v ,  , k , R are the density, SO3 mass fraction, SO3 diffusion flux, pressure, temperature, velocity, stress tensor, conductivity, and gas constant, respectively. Furthermore, S m is the mass source added into the continuum phase from the discrete phase, due to Ca(OH)2 sulfuration; SSO3 is the creation rate by addition from the dispersed phase; Smv


8

Page 9 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

is the external body force source arising from, for example, interaction with the discrete phase;

Srad is the energy source due to radiation, calculated by the P-1 radiation model, with absorption coefficients calculated by the weigh-sum-of-gray-gas model (WSGGM); and S h is the energy source due to the gas-solid reaction of Ca(OH)2 with SO3. The energy e is calculated as follows:

v2 2

(6)

h  Yj hj ,

(7)

e  h

p





j

where Y j and h j are the mass fraction and sensible enthalpy of species j respectively. The diffusion flux of species j ( J j ) is obtained according to Fick’s law: J j    D j , m Y j ,

(8)

where D j ,m is the mass diffusion coefficient for species j. 3.2 Discrete phase model The Ca(OH)2 sorbent behaviors are described as the discrete phase, which is solved by tracking a large number of particles through the calculated flow field. The trajectory of a discrete phase particle is calculated by integrating the force balance equation on the particle: d vp dt

 FD (v  v p ) 

g ( p   )

p

F ,

(9)

where v is the gas velocity, v p is the particle velocity,  is the gas density,  p is the particle density, F is an additional acceleration term, FD (v  u p ) is the drag force term, and


9

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

FD 

Page 10 of 30

18 CD Re .  p d p2 24

(10)

In the above,  is the molecular viscosity, d p is the particle diameter, and Re is the relative Reynolds number: Re 

d p v  vp 

.

(11)

Following the trajectory of a discrete phase particle, a heat and mass exchange also occurs between the discrete and gas phases. For mass balance consideration, the Ca(OH)2 mass depletion rate can be expressed by either the chemical reaction or diffusion rate:

d mp dt

 Ap M p k g  Cg  Cs   Ap M p kr  Cs  , n

(12)

where k g (Cg -Cs ) is the reactant gas diffusion rate from the bulk to sorbent particle surface, and kr (Cs )n is the apparent chemical reaction rate, with a reaction order of n . Moreover, M p is

the sorbent molecular weight, and the diffusion rate coefficient k g can be calculated as k g  Sh

According to Smith et al.,

32

Dg dp

.

(13)

the Sherwood number can be taken as 2 for the SO3 removal test,

where the relative velocity between the gas and particle is usually small. Here, Dg denotes the reactant gas diffusion coefficient. The apparent chemical reaction rate coefficient is generally obtained as follows: kr  A exp( E / RTp )(1  X ) ,

(14)

where the chemical reaction of the SO3 removal test can be expressed by global reactions, as

Ca(OH)2  SO3  CaSO4  H2O .


(15)

10

Page 11 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

In the above, X is the sorbent particle conversion, and X = 1 when the sorbent particle reacts completely. Assuming that the reaction order is 1, the SO3 removal reaction rate is calculated as a combination of the diffusion rate and apparent chemical reaction rate: d mp dt

 Ap M p

kr k g kr  k g

(16)

Cg .

The particle diameter is assumed to be constant, while the density decreases during reaction. The reactant gas concentration C g is often replaced by the reactant gas partial pressure pg . The particle temperature is calculated by considering the heat transfer of conduction, convection, radiation and reaction heat:

mpc p

d Tp dt

 Ap hcon Tg  Tp   Ap p Tr4  Tp4  

d mp dt

H sulf ,

(17)

where m p , Ap , c p ,  p are the particle mass, surface area, heat capacity, and emissivity, respectively, while Tp , Tg , Tr are the particle, gas, and gas radiation temperatures, respectively. Furthermore,



is the Stefan-Boltzmann constant, and H sulf is the reaction heat due to the

reaction of Ca(OH)2 with SO3. Moreover, hcon is the convection transfer coefficient, calculated from correlation, as follows: hcon 

g dp

 2.0  0.6 Re

1/2 g ,d p



Prg1/3 ,

(18)

where g is the gas thermal conductivity; Re g ,d p is the gas Reynolds number based on the particle diameter; and Prg is the gas Prandtl number. As the time step ( t ) for particle trajectory is small, an approximate solution is applied to calculate particle temperature changes within one time step ( t ), assuming that the particle temperature and mass do not change significantly. The integration from time t to time t  t can be expressed by


11

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 12 of 30

TP (t  t )   P  TP (t )   P  e 

P t

,

(19)

where

d mp H sulf  Ap p Tr4 dt hcon Ap  Ap p TP3

hcon ApTg 

P 

P 

Ap (h   p TP3 ) mp c p

.

(20)

(21)

3.3 Modeling approach and grids The above models were conducted using the commercial CFD software Fluent. In the CFD simulation, the Reynolds-averaged Navier-Stokes (RANS) model was used for turbulence calculation, while the realizable k   model was applied for turbulence closure. The calculation was based on the Eulerian-Lagrangian formulation for the continuous gas phase and dispersed particle phase. The particle motions were simulated by means of a stochastic trajectory model. In addition, the SIMPLE algorithm was applied for coupling of the velocity and pressure field. Governing equations were discretized by the first-order upwind scheme, and certain default under-relaxation factors were reduced in order to maintain effective convergence. Iteration calculations were carried out until the solution satisfied a pre-specified tolerance. Based on the model described above, the Ca(OH)2 sulfuration model was programmed into a user-defined function (UDF) file and implemented in Fluent, using the “Define_DPM_LAW” macro. The sulfuration reaction and heat transfer calculation of the particle were included in the UDF. A frame of the UDF is illustrated in Figure 4.


12

Page 13 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

Figure 4. Schematic of Ca(OH)2 sulfuration model implemented in Fluent.

A 3D grid of the experimental system is illustrated in Figure 5. Grid independence was conducted on three different numbers of cells. The predicted gas temperature and reactant gas concentration values at the exit were used as criteria. The 0.23 million cells were eventually adopted, as this number provides a similar prediction result with finer mesh and a smaller computational cost. The sorbent particles were divided into 10 groups with the same particle size, and 2420 particle trajectories were tracked in the simulation.

Figure 5. 3D grid of experimental system.

4 Results


13

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 14 of 30

4.1 Model validation In order to validate the proposed model, the SO3 removal process along the reaction distance was calculated under different temperatures and [Ca]/[S] values, using the above models. The simulated results are illustrated in Figure 6 (a) and compared with the experimental data in Figures 6 (b)—(d), where the lines are the calculated results and the dots are experimental data. As shown in Figure 6 (a), the SO3 concentration decreases due to reaction (15) along the reaction distance, and the mass of the sorbent particles increases accordingly. These results indicate that the SO3 removal process inside the EFR was reasonably simulated. It can be seen from Figures 6 (b)—(d) that the SO3 removal efficiency obtained using the model agrees reasonably strongly with the experimental data, which implies that the model developed in this work can effectively predict SO3 removal reaction performance. In the model, the chemical reaction rate is described in Arrhenius form, and the optimum values of the kinetic parameters A and E can be determined by minimizing the variance between experimental data and modeling results, including temperature and gas concentration. The obtained pre-factor A is 1.9 kg/(m2 ·s·Pa) and the activation energy E is 45 kJ/mol.

(a)


14

Page 15 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

(c)

(b)

(d) Figure 6. (a) Simulated results of removal process in reactor at 400 oC; (b)—(d) Comparison of experimental data and calculated results: (b) reaction temperature is 290 oC; (c) reaction temperature is 350 oC; (d) reaction temperature is 400 oC (lines are calculated results and dots are experimental data).

It can be observed from Figure 6 (a) that there is an obvious SO3 removal effect when Ca(OH)2 is injected into the reactor. Furthermore, the temperature has a significant influence on SO3 removal efficiency, which affects the design of the residence time. In order to cause the SO3 concentration to be lower than 5 ppm, as recommended in the literature,

25

the SO3 removal

efficiency should be higher than 65%. In this case, the residence time at 400 oC can be designed as lower than that at 350 oC or 290 oC in order to ensure SO3 removal efficiency of higher than


15

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 16 of 30

65% at the reactor outlet. Furthermore, temperature will affect the [Ca]/[S] selection. When the temperature is 400 oC, the molar ratio of 5 is large enough to achieve satisfactory SO3 removal efficiency; that is, a higher reaction temperature can save additional sorbents. 4.2 Effect of sorbent injection speed on SO3 removal efficiency The difference between the sorbent injection and gas mixture speeds will affect the mixing degree of the gas and sorbent particles flows. During the experiment, the sorbent injection speed can be changed by adjusting the flow rate of the gas carrying sorbent particles. Figure 7 illustrates the simulation results of SO3 removal efficiency under different sorbent injection speeds. It can be observed that the highest SO3 removal efficiency is achieved when the sorbent injection speed is lower than the flue gas speed. Furthermore, a higher sorbent injection speed will result in lower SO3 removal efficiency. The main reason is that a poor degree of mixing exists between the gas and particles when the sorbent injection speed is higher than the gas mixture speed, as shown in Figures 7 (b)—(c), while the amount of reactant gas diffusing to the particle unit surface area will decrease. Therefore, a lower sorbent injection speed is conducive to increasing SO3 removal efficiency, which can be carried out by decreasing the carrier gas flow rate or increasing the injector tube diameter.

(a)


16

Page 17 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

(b)

(c)

Figure 7. (a) Effect of sorbent injection speed on SO3 removal efficiency; (b) distribution of particle trajectories; and (c) distribution of gas velocity (temperature is 290 oC).

4.3 Effect of sorbent properties on SO3 removal efficiency The sorbent particle diameter significantly affects the mass transfer process. As indicated in equation (13), a smaller sorbent particle diameter corresponds to a larger reactant gas diffusion rate coefficient, which means that SO3 can contact the sorbent particle surface more easily. Furthermore, the total surface area will be larger for the same feeding amount. Consequently, the contact ratio between the SO3 and sorbent particle will be larger and the SO3 removal reaction rate will be higher. Therefore, it is important to analyze the effects of sorbent particle diameter on reaction performance. Figure 8(a) illustrates the simulation results of SO3 removal efficiency under different sorbent particle diameters, from 3 to 40 μm, and it can be seen that the sorbent particle diameter has a significant influence on performance. There is an obvious increasing trend in SO3 removal efficiency when decreasing the sorbent particle diameter. For example, when the diameter decreases from 10 to 3 μm, SO3 removal efficiency at the reactor outlet could increase from 55% to 99%. Therefore, sorbent with a particle diameter that is as small as possible is more suitable for removing SO3 and achieving a significantly higher SO3 removal efficiency. However, an excessively small sorbent will increase the sorbent preparation cost and result in sorbent feeding


17

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 18 of 30

problems. The sorbent reactivity will affect the sorbent and SO3 reaction rate. In the model, the sorbent reactivity can be changed by adjusting the pre-factor A. The effect of sorbent reactivity on SO3 removal efficiency under three different sorbent particle diameters (6, 20, and 40 μm) was investigated in this study, as illustrated in Figures 8(b)—(d). Extremely diverse reaction behaviors can be observed for the different sorbent particle diameters. When the sorbent particle diameter is 6 μm, SO3 removal efficiency at the reactor outlet can increase from 70% to 95% by increasing the pre-factor from 1 to 10, which means that sorbent reactivity can significantly affect SO3 removal efficiency. Here, the reaction process is controlled by a chemical reaction; however, when the pre-factor is larger than 50, a further increase will not significantly affect the reaction performance in the reactor, which can be attributed to the fact that the chemical reaction rate is sufficiently fast and the gas diffusion becomes the reaction controlling mechanism. When the sorbent particle diameter is 20 μm, SO3 removal efficiency at the reactor outlet can only increase to 30% by improving sorbent reactivity and only a slight change occurs in SO3 removal efficiency for different pre-factors, because a higher sorbent particle diameter results in a higher reactant gas diffusion resistance. Furthermore, when the sorbent particle diameter increases to 40 μm, the sorbent reactivity has almost no effect on SO3 removal efficiency, and the efficiency at the reactor outlet remains at only 10%, in which case there is no point in improving the sorbent reactivity. Based on the above analysis, the effect of sorbent reactivity on reaction performance is severely affected by the sorbent particle size value. In subsequent parts of this work, the controlling mechanism of the Ca(OH)2 reaction with SO3 will be investigated in detail by means of theoretical derivation.


18

Page 19 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

(a)

(b)

(c)

(d)

Figure 8. (a) Effect of sorbent particle diameter on SO3 removal efficiency; (b—d) effect of sorbent reactivity on SO3 removal efficiency under three different sorbent particle diameters: (b) 6 μm; (c) 20 μm; and (d) 40 μm (temperature is 290 oC).

5 Discussion Considering the small amount of SO3 in flue gas (10~60 ppm 13), competition exists between the external gas diffusion and gas-solid reaction. For the gas-solid reaction of Ca(OH)2 with SO3, the injection flow rate of the sorbent feed, M, is a constant if the SO3 concentration in the flue gas and [Ca]/[S] are specified. At the reactor inlet reaction zone, the overall reaction rate is calculated as follows:


19

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

rN

d mp dt

 N  Ap M p

Page 20 of 30

kr k g kr  k g

Cg ,

(22)

where r is the overall reaction rate and N denotes the total number of sorbent particles injected into the reactor per unit time. Then, N can be calculated as N

M 1  p d p3 6 ,

(23)

where  p is the particle density and d p is the particle diameter. Considering that the external surface area of each sorbent particle is Ap   d p2 , the overall reaction rate can be obtained as follows: r

kk 6M  M p r g Cg .  pd p kr  k g

(24)

Furthermore, considering that k g is a function of the particle diameter, as shown in equation (13), the reaction rate can be calculated as follows:

r

6M Mp  pd p

The relative magnitudes between k r

kr  Sh kr  Sh

and k g

Dg dp C . Dg g

(25)

dp can determine the rate controlling

mechanism. If k g is significantly smaller than k r , the overall reaction rate is simplified as

r

6 MM p Dg C g Sh 1  2. p dp

(26)

In this case, the overall reaction rate is completely controlled by the external gas diffusion, and this mechanism is known as Zone III.


20

Page 21 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

Similarly, if k r is significantly smaller than k g , the overall reaction rate is simplified as

r

6MM p Cg

p

 kr 

1 . dp

(27)

The overall reaction rate is completely controlled by a chemical reaction, and this mechanism is known as Zone I. If k r and k g are within the same magnitude range, the overall reaction rate is controlled by both the external gas diffusion and chemical reaction, and this mechanism is known as Zone II. The relation curve between the reaction rate and sorbent particle diameter can be obtained by equation (25), as shown in Figure 9. A summary of the parameter values is provided in Table 2. Considering that the diffusion coefficient of SO3, Dg, cannot be determined directly in relative reference, Dg is calculated by the following relationship with DO2 : 33

Dg DO2

M O2



M SO3

(28)

.

Table 2. Values of parameters used in Figure 9

M

p

Mp

Cg

Sh

Dg

A

E

kg/s

kg/m3

kg/mol

mol/m3

1

m2/s

m/s

J/mol

2.32 × 10-5

2.24 × 103

7.40 × 10-2

7.21 × 10-4

2

3.33 × 10-5

1.9

4.50 × 104


21

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 22 of 30

Figure 9. Relation curve between reaction rate and sorbent particle diameter (I: chemical reaction control; II: combination control; III: gas diffusion control).

It can be observed from Figure 9 that the sorbent diameter has a significant effect on the overall reaction rate, and a typical characteristic of three zones exists. A very interesting finding is that the overall reaction rate is linearly dependent on 1/ d p2 in zone III, as shown in eq. (26), indicating that external gas diffusion is the rate controlling step in zone III and the sorbent diameter is a dominant parameter, while the chemical reaction kinetics have almost no effect on the overall reaction rate in zone III. With the decrease in sorbent diameter, 1/ d p2 will increase, resulting in an increase in external gas diffusion, as illustrated in Figure 9. When the external gas diffusion is within the same magnitude range as the gas-solid reaction, the overall reaction is located in Zone II, in which case the gas-solid reaction competes with the external gas diffusion, as illustrated in Figure 9. Upon further decreasing the sorbent diameter, the external gas diffusion becomes significantly greater than the gas-solid reaction rate and the overall reaction is located in Zone I, which is controlled by the chemical kinetics of Ca(OH)2 with SO3. In practical applications, the sorbent diameter should be sufficiently small to eliminate the external diffusion, in order to remove SO3 efficiently within a very short residence time. Based on the above discussion, the controlling mechanism in the three zones shown in Figure 9 should be considered


22

Page 23 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

for practical DSI technology.

6 Conclusions A pilot-scale EFR system was designed and constructed in order to investigate SO3 removal using a dry sorbent injection of Ca(OH)2 particles. In this EFR system, SO3 was stably produced through catalytic oxidation of SO2, using a vanadium-based catalyst, and the controlled condensation method was used to measure the SO3 concentration in the flue gas. The experimental results demonstrated that increasing the temperature and [Ca]/[S] can improve SO3 removal performance. The experimental data in the EFR were interpreted with the CFD simulation; the SO3 removal reaction model was established based on the gas-solid reaction principle and implemented in the CFD frame by means of a user-defined function. The CFD model results fit the experimental information reasonably well. CFD simulation was applied to investigate the effect of certain main parameters on SO3 removal efficiency by means of a sensitivity analysis, including sorbent injection speed, particle diameter, and reactivity. The simulation results indicated that the sorbent particle size has the greatest influence on removal efficiency. When the sorbent particle diameter decreases from 10 to 3 μm, SO3 removal efficiency could increase from 55% to 99%. When the sorbent particle diameter is larger than 20 μm, almost no effect of sorbent reactivity on SO3 removal efficiency is observed. According to the theoretical derivation and discussion, the SO3 removal process by Ca(OH)2 is controlled by both the gas-solid reaction and gas external diffusion. For large sorbent particles, the gas diffusion is the rate controlling step; in this case, sorbent reactivity has little influence on SO3 removal reaction performance. Therefore, for practical applications, the sorbent particle size should be sufficiently small to reduce the gas diffusion resistance, which will further result in an obvious improvement in the sorbent reaction performance.


23

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 24 of 30

* Corresponding author. Mailing address: Department of Energy and Power Engineering, Tsinghua University, Beijing 100084, China Telephone: 86-10-62789955 Fax: 86-10-62770209 E-mail: [email protected]

Acknowledgments This research was supported by the National Key Research and Development Program of China (No. 2016YFB0600802-A), Sichuan Province—University Science and Technology Cooperation (No. 2017JZ0002), and the EU H2020 Project (No. 764816 CLEANKER).

Nomenclature A

pre-factor

Ap

particle surface area

cp

particle heat capacity

CD

drag force coefficient

Cg

concentration of reactant gas in bulk

Cs

concentration of reactant gas on sorbent particle surface

dp

particle diameter

Dg

diffusion coefficient of reactant gas

Dj,m

mass diffusion coefficient for species j

DO2

diffusion coefficient of O2

e

energy

E

activation energy


24

Page 25 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

⃗F

additional acceleration term

FD

drag force

h

sensible enthalpy

hcon

convection transfer coefficient

hj

sensible enthalpy of species j

Jj

diffusion flux of SO3

JSO3

diffusion flux of SO3

k

conductivity

kg

diffusion rate coefficient

kr

chemical reaction rate coefficient

mp

particle mass

M

injection flow rate of sorbent feed

Mp

molecular weight of sorbent

n

reaction order

N

total number of sorbent particles injected into reactor per unit time

p

gas pressure

pg

partial pressure of reactant gas

Prg

Prandtl number of gas

r

reaction rate

R

gas constant

Re

relative Reynolds number

Reg,dp

Reynolds number of gas based on particle diameter

Sh

energy source due to gas-solid reaction


25

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Sm

mass source added into continuum phase from discrete phase

Smv

external body force source

Srad

energy source due to radiation

SSO3

creation rate by addition from dispersed phase

Sh

Sherwood number

t

time

T

temperature

Tg

gas temperature

Tp

particle temperature

Tr

gas radiation temperature

v

gas velocity

vp

particle velocity

Yj

mass fraction of species j

YSO3

SO3 mass fraction

ΔHsulf

reaction heat

εp

particle emissivity

λg

gas thermal conductivity

μ

molecular viscosity

ρ

gas density

ρp

particle density

σ

Stefan-Boltzmann constant

τ̿

stress tensor

Page 26 of 30

References


26

Page 27 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

(1) Attar, A. Chemistry, thermodynamics and kinetics of reactions of sulphur in coal-gas reactions: A review. Fuel 1978, 57 (4), 201–212. (2) Raask, E. Sulphate capture in ash and boiler deposits in relation to SO2 emission. Prog. Energy Combust. Sci. 1982, 8 (4), 261–276. (3) Fleig, D.; Andersson, K.; Normann, F.; Johnsson, F. SO3 Formation under Oxyfuel Combustion Conditions. Ind. Eng. Chem. Res. 2011, 50 (14), 8505−8514. (4) Kamata, H.; Ohara, H.; Takahashi, K.; Yukimura, A.; Seo, Y. SO2 oxidation over the V2O5/TiO2 SCR catalyst, Catal. Lett. 2001, 73 (1), 79–83. (5) Schwaemmle, T.; Heidel, B.; Brechtel, K.; Scheffknecht, G. Study of the effect of newly developed mercury oxidation catalysts on the DeNOx-activity and SO2–SO3-conversion. Fuel 2012, 101, 179–186. (6) Schwammle, T.; Bertsche, F.; Hartung, A.; Brandenstein, J.; Heidel, B.; Scheffknecht, G. Influence of geometrical parameters of honeycomb commercial SCR-DeNOx-catalysts on DeNOx-activity, mercury oxidation and SO2/SO3-conversion. Chem. Eng. J. 2013, 222 (8), 274−281. (7) Svachula, J.; Alemany, L. J.; Ferlazzo, N.; Forzatti, P.; Tronconi, E.; Bregani, F. Oxidation of SO2 to SO3 over Honeycomb DeNoxing Catalysts. Ind. Eng. Chem. Res. 1993, 32 (5), 826−834. (8) Tronconi, E.; Cavanna, A.; Orsenigo, C.; Forzatti, P. Transient Kinetics of SO2 Oxidation over SCR-DeNOx Monolith Catalysts. Ind. Eng. Chem. Res. 1999, 38 (7), 2593−2598. (9) Brachert, L.; Kochenburger, T.; Schaber, K. Facing the Sulfuric Acid Aerosol Problem in Flue Gas Cleaning: Pilot Plant Experiments and Simulation, Aerosol Sci. Technol. 2013, 47 (10) 1083–1091. (10) Otsuka, N. Effects of fuel impurities on the fireside corrosion of boiler tubes in advanced


27

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 28 of 30

power generating systems—a thermodynamic calculation of deposit chemistry. Corros. Sci. 2002, 44 (2), 265–283. (11) Xiang, B. X.; Zhang, M.; Yang, H. R.; Lu, J. F. Prediction of Acid Dew Point in Flue Gas of Boilers Burning Fossil Fuels. Energy Fuels 2016, 30 (4), 3365−3373. (12) Xiang, B. X.; Tang, B.; Wu, Y. X.; Yang, H. R.; Zhang, M.; Lu, J. F. Predicting acid dew point with a semi-empirical model. Appl. Therm. Eng. 2016, 106 (5), 992−1001. (13) Krishnakumar, B.; Niksa, S. Predicting the impact of SO3 on mercury removal by carbon sorbents. Proc. Combust. Inst. 2011, 33 (2), 2779−2785. (14) Xiang, B. X.; Shen, W. F.; Zhang, M.; Yang, H. R.; Lu, J. F. Effects of different factors on sulfur trioxide formations in a coal-fired circulating fluidized bed boiler. Chem. Eng. Sci. 2017, 172, 262−277. (15) Sinanis, S.; Wix, A.; Ana, L.; Schaber, K. Characterization of sulphuric acid and ammonium sulphate aerosols in wet flue gas cleaning processes. Chem. Eng. Process. 2008, 47 (1), 22–30. (16) Buckley, W. P.; Altshuler, B. Acid mist causes problems for FGD systems. Power Eng. 2002, 106 (11), 132–136. (17) Xiang, B. X.; Zhang, M.; Wu, Y. X.; Yang, H. R.; Zhang, H.; Lu, J. F. Experimental and Modeling Studies on Sulfur Trioxide of Flue Gas in a Coal-Fired Boiler. Energy Fuels 2017, 31 (6), 6284–6297. (18) Wang, Z. Q.; Huan, Q. C.; Qi, C. L; Zhang, L. Q.; Cui, L.; Xu, X. R.; Ma, C. Y. Study on the Removal of Coal Smoke SO3 with CaO. Energy Procedia 2012, 14, 1911–1917. (19) Adams, B.; Senior, C. Curbing the blue plume: SO3 formation and mitigation. Power 2006, 150 (4), 39–41. (20) Muzio L. J.; Offen, G. R. Assessment of Dry Sorbent Emission Control Technologies: Part I.


28

Page 29 of 30 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Energy & Fuels

Fundamental Process. JAPCA 1987, 37 (5), 642–654. (21) Often, G. R.; Mcelroy, M. W.; Muzio, L. J. Assessment of Dry Sorbent Emission Control Technologies: Part II. Applications. JAPCA 1987, 37 (8), 968–980. (22) Yelverton, T. L. B.; Nash, D. G.; Brown, J. E.; Singer, C. F.; Ryan, J. V.; Kariher, P. H. Dry sorbent injection of trona to control acid gases from a pilot-scale coal-fired combustion facility. AIMS Environ. Sci. 2016, 3 (1), 45–57. (23) Thibault, J. D.; Steward, F. R.; Ruthven, D. M. The Kinetics of Absorption of SO3 in Calcium and Magnesium Oxides. Can. J. Chem. Eng. 1982, 60 (12), 796–801. (24) Chen, P.; Wang, Z. Q.; Chang, J. C.; Ma, C. Y. Experimental Study of the Reactivity of Ca-Based Matters with SO3. Power and Energy Engineering Conference, 2011; pp 1–4. (25) Moretti, A. L.; Triscori, R. J.; Ritzenthaler D. P. A System Approach to SO3 Mitigation. Combined Power Plant Air Pollutant Control Mega Symposium, USA, 2006; pp 1–7. (26) Huang, R. T.; Yu, R.; Wu, H.; Pan, D. P.; Zhang, Y. P.; Yang, L. J. Investigation on the removal of SO3 in ammonia-based WFGD system. Chem. Eng. J. 2016, 289, 537–543. (27) Ueda, Y.; Nagayasu, H.; Hamaguchi, R.; Miyake, K.; Matsuura, K.; Nagata, C. SO3 Removal System for Flue Gas in Plants Firing High-sulfur Residual Fuels. Mitsubishi Heavy Industries Technical Review 2012, 49 (4), 6–12. (28) Gray, S. M.; Jarvis, J. B.; Kosler, S. W. Combined Mercury and SO3 Removal Using SBS Injection. Power 2014, 158 (7), 22–26. (29) Kocaefe, D.; Karman, D.; Steward, F. R. Comparison of the Sulfation Rates of Calcium, Magnesium and Zinc Oxides with SO2 and SO3. Can. J. Chem. Eng. 1985, 63 (6), 971–977. (30) Steward, F. R.; Karman, D.; Kocaefe, D. A Comparison of the Reactivity of Various Metal Oxides with SO3. Can. J. Chem. Eng. 1987, 65 (2), 342–344.


29

Energy & Fuels 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

Page 30 of 30

(31) Moser, R. E. SO3's impacts on plant O & M: part I. Power 2006, 150 (8), 40–40. (32) Smith, I. W. The combustion rates of coal chars: A review. Proceedings of the 19th International Symposium on Combustion, The Combustion Institute, Pittsburgh, PA, 1982; pp 1045-1065. (33) Graham, T. On the law of the diffusion of gases. J. Membr. Sci. 1995, 100 (1), 17–21.


30

Removal from Flue Gas with Ca(OH) - American Chemical Society

Recommend Documents