Modeling of Mercuric Chloride Removal by CuCl2-Impregnated

Modeling of Mercuric Chloride Removal by CuCl2-Impregnated Activated Carbon ... (1) It is estimated that over 300 new fabric filters (FFs) and even mo...
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Modeling of Mercuric Chloride Removal by CuCl2‑Impregnated Activated Carbon Sorbent in a Fabric Filter Xin Li and Joo-Youp Lee* Chemical Engineering Program, Department of Biomedical, Chemical, and Environmental Engineering, University of Cincinnati, Cincinnati, Ohio 45221-0012, United States ABSTRACT: A model for the physical adsorption of mercuric chloride (HgCl2) onto raw activated carbon (AC) and 4 and 10% CuCl2-impregnated activated carbon (CuCl2−AC) sorbents in a fabric filter was studied on the basis of the adsorption equilibrium and kinetics. Sorbent loading, particle size, filtration time, and CuCl2 loading onto AC were found to be the major factors determining a HgCl2 removal efficiency. Although high sorbent loading and small sorbent particle size result in high HgCl2 removal efficiencies, the inlet HgCl2 concentration and superficial gas velocity were found to have little impact on HgCl2 removal. Our operation case study result shows that sorbent injection for a short duration at a high loading followed by the discontinuation of the injection until a cleaning cycle can significantly save a total sorbent amount with high sorbent utilization, while the same amount of HgCl2 is removed. The adsorption of HgCl2 onto raw AC and CuCl2−AC in the filter cake was found to be primarily governed by the adsorption capacity difference in the Langmuir adsorption kinetic expression. The study results demonstrate how fundamental adsorption equilibrium and kinetics can be used to design a sorbent and predict its performance in a fabric filter. A comparison of the current model predictions was also made with pilot-scale field data available in the literature.

1. INTRODUCTION Because of the impending Mercury and Air Toxics Standards (MATS) rule set forth by the United States Environmental Protection Agency (U.S. EPA), many U.S. coal-fired power plants are slated to be equipped with necessary air pollution control equipment in the years to come.1 It is estimated that over 300 new fabric filters (FFs) and even more electrostatic precipitators (ESPs) are expected to be installed for the capture of mercury and fine particulate.2 Simulations of in-flight capture of oxidized mercury (Hg2+) vapor by raw AC sorbents in flue gas stream were reported to be insignificant.3−5 A field test result obtained during the U.S. Department of Energy’s (DOE’s) Phase II Mercury Control Technology Field Testing Program also suggests that in-flight capture of mercury vapor in the ductwork does not appear to be significant with short residence times.6 Mercury removal efficiency in a FF is significantly higher than that in an ESP for both bituminous and sub-bituminous coals because of increased gas−sorbent contact in the filter cake accumulated inside the filter bags.7,8 Therefore, it is a filter cake comprising injected activated carbon (AC) sorbents that primarily adsorbs substantial mercury vapor. The filter cake acts as a fixed-bed system with growing thickness. Previous studies have shown that raw AC cannot physically adsorb elemental mercury (Hg0) vapor in a post-combustion temperature window.3,8,9 Among chemically promoted AC sorbents, halogenated ACs have been reported to significantly enhance Hg0 vapor adsorption.8,9 These chemically treated AC sorbents have demonstrated the capability to adsorb both elemental and oxidized forms of mercury vapor from coal combustion flue gases.8,10,11 However, it is difficult to find mercury adsorption modeling studies for the prediction of elemental and oxidized mercury vapor removal by chemical and physical sorbents in a FF primarily because of the difficulty in finding the equilibrium and kinetic data associated with reaction © 2013 American Chemical Society

and adsorption mechanisms between mercury vapor and a sorbent. A previous study about the adsorption kinetics of HgCl2 onto raw and sodium sulfide (Na2S)-impregnated AC in a lab-scale fixed bed showed that impregnated Na2S impeded the adsorption kinetics of HgCl2 in comparison to the adsorption onto raw AC.12 Other modeling studies of HgCl2 adsorption onto several sorbents using adsorption kinetic parameters and physical properties, including adsorption/ desorption kinetic constants, maximum adsorption capacity, particle density/porosity, and pore diameter, are reported in the literature.13,14 The models considered external mass transfer from the bulk gas phase to the external surface of the AC particle and intraparticle pore diffusion within the particle. In our previous work, cupric chloride (CuCl2)-impregnated AC (CuCl2−AC) showed the capability to adsorb both Hg0 and Hg2+ species.15,16 It was found that Hg0 vapor first reacts with CuCl2 impregnated onto the carbon surface, and resultant HgCl2 is desorbed from the CuCl2 site and subsequently readsorbed onto CuCl2-free carbon sites by physical adsorption. Therefore, it is essential to separately determine the chemical reaction and physical adsorption kinetics for the prediction of Hg0 vapor removal by CuCl2−AC for sorbent injection. Our recent study about the adsorption equilibrium and kinetics for HgCl2 adsorption onto the CuCl2−AC sorbent showed that CuCl2 loading does not decrease the HgCl2 adsorption kinetics and increases the binding energy of HgCl2 adsorption.16 On the basis of the equilibrium and kinetic study results, in this study, we have extended our model to predict HgCl2 vapor removal in the filter cake by CuCl2−AC sorbent in comparison to raw AC sorbent. Received: August 31, 2013 Revised: November 6, 2013 Published: November 12, 2013 7654

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Figure 1. Schematic of HgCl2 adsorption by AC sorbent in a FF.

2. KINETIC MODEL OF HGCL2 VAPOR ADSORPTION ON AC

rA = ρp

∂q = ρp (k1(qmax − q)C − k 2q) ∂t

(1)

At equilibrium (rA = 0), the rate equation leads to the Langmuir isotherm.

2.1. Model Assumptions. The adsorption of HgCl2 on raw AC (Norit DARCO-HG) and CuCl2−AC was identified as physical adsorption.12,16 The Langmuir theory has been adopted for this physical adsorption study because of its simple expression and success in correlating experimental adsorption data.3,12,13,16 Because mercury concentrations in flue gases are very low [usually on the order of a few parts per billion by volume (ppbv)], significant external and internal mass-transfer resistances exist for sorbent injection. The HgCl2 adsorption onto carbon sorbents consists of a series of three steps: (1) external mass transfer from the bulk gas phase to the external surface of a sorbent particle, (2) intraparticle mass transfer from the external carbon surface to the interior of the particle through particle pores, and (3) surface adsorption on the intraparticle surface area of a sorbent. In this study, an adsorption kinetic model for a FF system was developed on the basis of the following additional assumptions: (1) Filter cake contains fly ash and injected AC sorbent, in which only the injected AC sorbent can adsorb HgCl2 vapor. Fly ash and AC sorbent have similar packing properties and densities. This will allow for two assumptions of a constant bed porosity and a replacement of a volume fraction by a mass fraction used in a HgCl2 mass balance inside the filter cake used in eq 4. (2) The gas passing through the filter cake travels in plug flow. (3) AC particles are spherical, and all of the particles have the same size and are uniformly dispersed in the filter cake. (4) The heat effect associated with HgCl2 adsorption is negligible because of its trace level concentrations. Therefore, the temperature of a sorbent in the system is assumed to be constant and uniform (i.e., isothermal). (5) The pressure drop across the filter cake is negligible, and thus, the gas velocity is constant along the filter cake. The pressure drop was estimated to be ∼1 kPa based on Darcy’s equation, and thus, the assumption was found to be reasonable. (6) The axial diffusion is negligible. 2.2. Model Equations. 2.2.1. Intrinsic Adsorption Rate Expression. The Langmuir theory gives the following HgCl2 net adsorption rate expression in eq 1:

q* = qmax

KC* 1 + KC*

(2)

The two parameters of qmax and K in the Langmuir isotherm are used for the Langmuir kinetic expression in eq 1. 2.2.2. Mass Balance Inside Sorbent Particles. A HgCl2 mass balance within the spherical sorbent particle is described in eq 3. A schematic for HgCl2 adsorption in a FF is shown in Figure 1.

εp

∂C ∂C ⎞ 1 ∂⎛ = 2 ⎜r 2De ⎟ − rA ⎝ ∂t ∂r ⎠ r ∂r

(3)

The local HgCl2 net adsorption rate expression in eq 1 is incorporated into the above local HgCl2 mass balance for the calculation of the local HgCl2 concentration (C). 2.2.3. Mass Balance in Filter Cake. A shell mass balance approach was used to describe the adsorption of HgCl2 inside the filter cake in the axial z direction (i.e., gas flow direction perpendicular to the filter cake plane), as shown in eq 4. The adsorption of HgCl2 was assumed to occur only inside the sorbent particles. An aforementioned assumption of similar densities for sorbent and fly ash can replace a volume fraction of an AC sorbent occupied in the filter cake consisting of the AC sorbent and fly ash with a mass fraction below

εb

mAC ∂C B ∂C + u B = − rA,obs(1 − εb) mash + mAC ∂t ∂z

(4)

in which

rA,obs =

3 ∂C De R p ∂r

(z , R p , t )

(5)

The observed adsorption rate expression at a sorbent radius of Rp was obtained by taking a volume average of the local HgCl2 adsorption rate inside the pores over the entire sorbent particle.16 The local HgCl2 7655

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Table 1. Physical Properties of Three Sorbents

a

sorbent

particle density, ρp (g/cm3)

BET surface area (m2/g)

pore volume, vp (cm3/g)

particle porosity, εp

dporea (nm)

DARCO-HG 4 wt % CuCl2−AC 10 wt % CuCl2−AC

0.950 0.980 1.04

476 414 340

0.53 0.49 0.39

0.50 0.48 0.41

6.2 6.3 6.4

dpore = average pore diameter of the sorbent.

Table 2. Langmuir Parameters and Adsorption Kinetic Constant Determined for Three Different Sorbents16 sorbent

qmax (g of HgCl2/g of sorbent)

K (m3/g of HgCl2)

k1 (m3 g−1 s−1)

reference

0.055 0.024 0.010

2.1 × 10 5.0 × 104 26.0 × 104

0.2 0.3 0.5

16 16 16

DARCO-HG 4 wt % CuCl2−AC 10 wt % CuCl2−AC

4

mass flux at the sorbent radius Rp is coupled with the bulk gas-phase HgCl2 concentration (CB) in eq 4 via eq 5. 2.2.4. Initial Conditions (ICs) and Boundary Conditions (BCs). A previous modeling study indicated that the in-flight capture of HgCl2 vapor between a sorbent injection location and a particulate matter control device (i.e., ESP or FF) would be insignificant with short gas residence times of a few seconds and reasonable sorbent loadings.3 Therefore, in this study, the in-flight capture of HgCl2 in the ductwork was assumed to be zero, which led to the following initial and boundary conditions:

q(z , r , t = 0) = 0;

IC

C B(z , t = 0) = 0; ∂q ∂r

BC ∂C ∂r

(z , r = 0, t )

= (z , r = R p , t )

kg De

∂C ∂r

∂γ ∂ξ

= 0;

q ; qmax

ξ≡

r ; Rp

λ≡

L ∂φ = F [k1(1 − φ)C Binγ − k 2φ] ∂τ u

L FDe

∂τ

z ; LF (8)

(9)

ρp qmax ξ 2R p2 ∂ ⎛ 2 ∂γ ⎞ [k1(1 − φ)C Binγ ⎜ξ ⎟− ∂ξ ⎝ ∂ξ ⎠ DeC Bin − k 2φ]

3(1 − εb)DeL F R p2u

mAC ∂γ mash + mAC ∂ξ

(λ ,1, τ )

∂φ ∂ξ

γ(λ , ξ , τ = 0) = 0; (12)

= 0; (λ , ξ = 0, τ )

= (λ , ξ = 1, τ )

k gR p De

(11)

∂γ ∂ξ

= 0; (λ , ξ = 0, τ )

[γB(λ , τ ) − γ(λ , ξ = 1, τ )]; (13)

3. RESULTS AND DISCUSSION A set of the above coupled partial differential equations (PDEs) (eqs 9−11) were solved using the finite element method under the platform of COMSOL Multiphysics (version 4.3a).17 A normalized HgCl2 concentration profile in the bulk gas phase [γB(λ, τ)] within the filter cake is shown in Figure 2 for 4% CuCl2−AC with a 20 μm particle diameter, a 0.1 g/m3 [=6.2 lb/million actual cubic feet (MMacf)] sorbent injection loading, and a 10 ppbv inlet HgCl2 concentration at 140 °C. The bulk HgCl2 concentration (γB) decreases along the filter cake thickness, and the outlet HgCl2 concentration on the FF surface [γB(λ = 0, τ)] decreases with an increase in time (i.e., growing filter thickness). The result shows that ∼90% removal of the inlet HgCl2 concentration is achievable after ∼70 min. The removal of HgCl2 obtained from 4% CuCl2−AC with a 20 μm sorbent diameter, a 0.1 g/m3 (=6.2 lb/MMacf) sorbent injection loading, and a 10 ppbv inlet HgCl2 concentration was selected as a base case simulation. A number of factors can affect the HgCl2 removal efficiency in the filter cake during sorbent injection. In this study, sorbent loading (mAC), sorbent particle diameter (Dp), inlet HgCl2 concentration (CBin), and

in which LF is the final cake thickness formed at a final filtration time (tF) before the cleaning of a FF.

=

=−

∂λ

2.3. Parameters in Partial Differential Equations. The model has been applied to DARCO-HG AC, 4 wt %, and 10 wt % CuCl2−AC sorbents. The adsorptive and physical properties of the sorbents are listed in Table 1. A bed porosity (εb) value of 0.46 was used in the model based on the bulk and particle density values for the AC sorbent. The determination of the effective pore diffusion coefficient (De) and the gas-phase mass-transfer coefficient (kg) is described in our previous study.16 In our previous work, the Langmuir parameters and intrinsic kinetic constants were determined from our fixed-bed tests by taking into account the adsorption kinetics, equilibrium, and internal and external mass transfer.16 These parameters are listed in Table 2 and were used in this study to predict the removal of HgCl2 vapor in a FF system.

(z , r = 0, t )

where L(t) is the growing filter cake thickness in m expressed by L = ((mash + mAC)u/ρb)t and ρb is the filter cake density. The filter cake thickness increases with respect to time. Thus, the inlet position where flue gas starts to enter the growing filter cake at z = L(t) is a moving boundary, and HgCl2 vapor leaves the FF at z = 0. CinB is the HgCl2 concentration in the bulk gas phase entering the filter cake, and kg is the gas-phase mass-transfer coefficient in m/s. The profiles of q(z, r, t), C(z, r, t), and CB(z, t) can be obtained by solving the coupled partial differential equations of eqs 1, 3, and 4 with the above initial and boundary conditions. 2.2.5. Normalization. The above equations along with the initial and boundary conditions have been normalized below by means of the following dimensionless variables:

εpuξ 2R p2 ∂γ

∂γB

γB(λ = L(τ )/L F , τ ) = 1

(7)

φ≡

+

φ(λ , ξ , τ = 0) = 0;

BC

(6)

C B(z = L , t ) = C Bin

C C ; γB ≡ inB ; C Bin CB u τ≡ t LF

∂τ

γB(λ , τ = 0) = 0

[C B(z , t ) − C(z , R p , t )];

γ≡

∂γB

IC

C(z , r , t = 0) = 0;

L(t = 0) = 0

= 0;

εb

(10) 7656

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to evaluate the performance of HgCl2 capture in the filter cake by sorbent injection. One is the normalized outlet HgCl2 in concentration [γB(λ = 0, τ) = Cout B /CB ], indicating a HgCl2 emission level at the FF outlet. The other is the normalized average HgCl2 uptake (qavg/qmax) by sorbent, representing the sorbent utilization, and qavg was calculated by t

qavg (t ) =

∫0 (CBin − CBout(t ))dt mACt

(14)

It is worthwhile to mention that HgCl2 removal in a FF system is not dependent upon a superficial gas velocity (u) and a fly ash concentration in flue gas (mash). This can be explained by eq 11. Substituting LF = ((mash + mAC)u/ρb)tF into eq 11 would give

Figure 2. HgCl2 concentration in the bulk gas phase of the filter cake [4% CuCl2−AC; CinB , 10 ppbv; sorbent loading, 0.1 g/m3 (=6.2 lb/ MMacf); Dp, 20 μm; and T, 140 °C].

εb

different sorbents (i.e., raw AC and CuCl2−AC) were examined for HgCl2 removal in a FF system. Two parameters were used

∂γB ∂τ

+

∂γB ∂λ

=−

3(1 − εb)Det F mAC ∂γ ρb ∂ξ R p2

(λ ,1, τ )

(15)

Figure 3. Effects of the sorbent injection loading on HgCl2 removal, for (a and b) continuous sorbent injection and (c and d) discontinuous sorbent injection. Y axis: (a and c) outlet HgCl2 concentration and (b and d) average HgCl2 uptake (4% CuCl2−AC; CinB , 10 ppbv; Dp, 20 μm; and T, 140 °C). 7657

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Figure 4. Effects of the sorbent particle size on HgCl2 removal. Y axis: (a) outlet HgCl2 concentration and (b) average HgCl2 uptake (4% CuCl2− AC; CinB , 10 ppbv; sorbent load, 0.1 g/m3; and T = 140 °C).

which is independent of u and mash. In reality, a superficial gas velocity influences the mass-transfer coefficient in the particle Sherwood number. However, the Sherwood number approaches the limiting value of 2 under typical gas/cloth ratios on the order of 1−4 cm/s in a FF.18 Thus, the influence of a gas velocity within the filter cake can also be ignored. Although different superficial gas velocities and fly ash loadings in flue gas do not influence HgCl2 removal in a FF at a specific filtration time, the pressure drop across the filter cake increases with an increase in the gas/cloth ratio or fly ash loading . On one hand, filter cake provides better contact between flue gas and sorbent. On the other hand, filter cake should be cyclically cleaned to maintain a reasonable pressure drop in a FF system. For typical reverse-gas cleaning, the pressure drop ranges from 0.9 to 2.2 kPa with the gas velocities of 0.8−1.2 cm/s.19 An allowable pressure drop determines a final filtration time, which is reported to range from 10 to 90 min in typical coal-fired power plants.19,20 Thus, all simulations were run up to 90 min in this study. 3.1. Effects of Sorbent Loading on HgCl2 Removal. Figure 3a shows the FF outlet HgCl2 concentration as a function of the filtration time using 4% CuCl2−AC with different loadings ranging from 0.01 g/m3 (=0.62 lb/MMacf) to 0.2 g/m3 (=12 lb/MMacf). According to the field test results obtained during the DOE’s Phase II Mercury Control Technology Field Testing Program, most plants do not seem to accommodate a sorbent injection rate greater than 0.16 g/m3 (=10 lb/MMacf).8 An inlet HgCl2 concentration and a sorbent particle diameter are 10 ppbv and 20 μm, respectively. It is clearly seen that sorbent loading has a significant impact on the HgCl2 emission level. The HgCl2 concentration at the outlet of a FF decreases with the filtration time for all four sorbent loadings as a sorbent accumulates on the filter cake and increases HgCl2 adsorption. However, the sorbent utilization efficiency increases with a decrease in sorbent loading with less than 7% of a maximum HgCl2 adsorption capacity of 0.024 g of HgCl2/g of 4% CuCl2−AC at the lowest sorbent loading of 0.01 g/m3, as shown in Figure 3b. This result indicates that most adsorption capacity of the sorbent is not used even with 90 min FF operation. When the MATS mandates >90% mercury emissions control, the simulation result shows that it will take more

than 35 min to start to remove >90% of an inlet 10 ppbv HgCl2 concentration even with 0.2 g/m3 sorbent loading of 4% CuCl2−AC. This is primarily because a certain amount of a sorbent is required to accumulate on the filter cake to achieve a desired removal efficiency (e.g., 90%). Therefore, the following operation study was carried out by injecting 4% CuCl2−AC at sorbent loadings between 0.1 and 2 g/m3 (i.e., 10 times higher than those used in the above continuous injection study) for the first 9 min (instead of 90 min) and discontinuing the injection until 90 min. This allows for the same amount of the sorbent to be injected over the 90 min filtration time until a cleaning cycle time. However, the discontinuous injection operation is expected to be able to achieve higher HgCl2 removal over the filtration time than the continuous one. Figure 3c shows a simulated operation case study result obtained from sorbent injection loadings between 0.1 and 2 g/m3. It clearly demonstrates that 9 min of injection at 1 g/m3 loading can achieve an average of 92% HgCl2 removal efficiency over a 90 min cleaning cycle time, while 90 min continuous injection at 0.1 g/m3 loading with the same total amount of the sorbent can achieve an average of 67% HgCl2 removal efficiency during the same 90 min cleaning cycle time. From the perspective of sorbent utilization, the continuous injection at 0.1 g/m3 loading over 90 min and the discontinuous injection at 1 g/m3 loading during the first 9 min give ∼3.1 and ∼4.3% sorbent saturation, respectively, as shown in panels b and d of Figure 3. A discontinuous short-duration injection method at a high sorbent loading will benefit high HgCl2 removal over the FF operation. 3.2. Effects of the Sorbent Particle Size on HgCl2 Removal. The simulation results show that the sorbent particle size is also an important factor affecting HgCl2 removal. Figure 4 represents the outlet HgCl2 concentration as a function of the filtration time in terms of different sorbent particle diameters ranging from 10 to 40 μm. The simulation was run for 4% CuCl2−AC with a 0.1 g/m3 loading and a 10 ppbv inlet HgCl2 concentration. The smaller the sorbent particle size, the lower the outlet HgCl2 concentration. For example, the simulation results estimated ∼40 and ∼70 min for 10 and 20 μm sorbents, respectively, until they started to remove 90% HgCl2. An average particle size for the AC used in this study (Norit DARCO-HG) was ∼15−20 μm. In our previous study, large 7658

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Figure 5. Effects of the inlet HgCl2 concentration on removal. Y axis: (a) outlet HgCl2 concentration and (b) average HgCl2 uptake (4% CuCl2−AC; sorbent loading, 0.1 g/m3; Dp, 20 μm; and T, 140 °C).

Figure 6. HgCl2 capture in terms of different sorbents, for (a and b) continuous sorbent injection at 0.1 g/m3 and (c and d) discontinuous sorbent injection, at 1 g/m3 for the first 9 min. Y axis: (a and c) outlet HgCl2 concentration and (b and d) average HgCl2 uptake (CinB , 10 ppbv; Dp, 20 μm; and T, 140 °C).

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Figure 7. Mercury removal efficiency as a function of the sorbent injection load at different filtration times at T = 140 °C: (a) continuous sorbent injection, filtration time = 10 min; (b) continuous sorbent injection, filtration time = 20 min; (c) continuous sorbent injection, filtration time = 50 min; (d) continuous sorbent injection, filtration time = 90 min; and (e) discontinuous sorbent injection, total filtration time = 90 min, with injection during the first 9 min.

respect to filtration time. The result obtained from 4% CuCl2− AC with a 20 μm diameter and 0.1 g/m3 loading shows that, within a range of 1−10 ppbv, the inlet concentration has little effect on the HgCl2 emission level, as shown in Figure 5a. It indicates that a difference in the driving force of the inlet HgCl2 concentration within the 1−10 ppbv concentration range is not significant for HgCl2 adsorption. However, an average mercury

particles showed a broad mass-transfer zone because of large intraparticle diffusional resistance, and small particles have a desired narrow mass-transfer zone that makes the sorbent more efficient.16 3.3. Effects of the Inlet HgCl2 Concentration on HgCl2 Removal. The effect of the inlet HgCl2 concentration on mercury removal was evaluated at the outlet of a FF with 7660

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Figure 8. Comparison of two model predictions to pilot-scale data (Dp = 20 μm).

concentrations fell rapidly within the first 9 min (down to 1, 5, and 13% of the inlet concentration for raw AC and 4 and 10% CuCl2−AC, respectively), and 95, 92, and 83% average removal efficiencies were obtained for raw AC and 4 and 10% CuCl2− AC, respectively. This enhanced removal is primarily derived from the early buildup of the sorbents in the filter cake. It is worthwhile to reiterate that CuCl2−AC is designed to capture Hg0 vapor and is used here to separately study the physical adsorption of oxidized mercury, HgCl2, formed as a result of the reaction between Hg0 vapor and CuCl2, while raw AC cannot capture Hg0 vapor at 140 °C. The result shows that a low CuCl2 loading, such as 4%, can achieve >90% HgCl2 removal at the same injection amount of raw AC for 9 min. This suggests that a CuCl2 loading for the CuCl2−AC sorbent can be determined depending upon the speciation of Hg0 and Hg2+ in flue gas. 3.5. Effects of Continuous and Discontinuous Injection on HgCl2 Removal Efficiency. The average mercury removal efficiency was calculated by taking an average of HgCl2 removal efficiencies during one cycle of filtration time, as shown in eq 16.

uptake for a high inlet concentration is greater than that for a low inlet gas mercury concentration, as shown in Figure 5b, indicating a higher sorbent utilization efficiency. Similar results were obtained for the sorbent with a range of 0.01−0.1 g/m3 sorbent loadings. 3.4. Effects of CuCl2 Loading on HgCl2 Removal. Figure 6 shows the outlet HgCl2 concentrations as a function of filtration time for three different sorbents of raw AC and 4 and 10% CuCl2−ACs. The HgCl2 adsorption capacities for CuCl2− AC are less than that for raw AC because HgCl2 is physically adsorbed onto CuCl2-free carbon site.21 It was also found that the adsorption kinetic constant (k1) increases with an increase in CuCl2 loading (i.e., 0.2, 0.3, and 0.5 for 0, 4, and 10% CuCl2−ACs, respectively) and the desorption kinetic constant (k2) decreases with an increase in CuCl2 loading because of an increase in the adsorption equilibrium constant (K).16 From Figure 6a, despite an increase in the adsorption kinetic constant (k1), the HgCl2 removal efficiencies for CuCl2−ACs are not predicted to be as high as that for raw AC. It indicates that the adsorption driving force of (qmax − q) in eq 1 dominates the kinetics for the physical adsorption of HgCl2 onto AC. Figure 6b also shows that the HgCl2 adsorption capacities for the three sorbents are still far less than their maximum adsorption capacities. To examine the effectiveness of a sorbent injection method of high loading followed by discontinuation for the three sorbents, the same operation case study shown in Figure 3 was run at a 1 g/m3 loading during the first 9 min. The results are reported in panels c and d of Figure 6. The outlet HgCl2

t

∫0 F (1 − γBout(t ))dt tF

(16)

Figure 7 shows average HgCl2 removal efficiencies as a function of sorbent loading at different filtration times. Filtration time and sorbent loading are the most important factors affecting the mercury removal efficiency. To obtain an average 90% HgCl2 7661

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removal efficiency after 90 min, the sorbent loadings of ∼0.23, 0.35, and 0.50 g/m3 would be required for raw AC and 4 and 10% CuCl2−AC, respectively, as shown in Figure 7d. These sorbent loadings for continuous injection are too high for most plants. However, if a discontinuous operation is applied, for the same 90% HgCl2 removal efficiency and 90 min filtration time, the required sorbent loadings are predicted to be 0.07, 0.09, and 0.13 g/m3 for raw AC and 4 and 10% CuCl2−AC, respectively, as shown in Figure 7e. As a result, more than 70% sorbent would be saved, while the same average 90% HgCl2 removal can be achieved during a 90 min cycle time. Between the two major FF cleaning methods, reverse-air cleaning usually has a cleaning cycle time of ∼30−90 min, while pulse-jet cleaning has to be cleaned approximately every 10 min.7 Therefore, to obtain a high mercury removal efficiency with a low sorbent injection loading, reverse-air cleaning is preferred over pulse-jet cleaning. 3.6. Comparison of Model Predictions to Literature Data. The model predictions made for HgCl2 removal by raw AC in the FF have been compared to pilot-scale data published in the literature.4,5,14 Flora et al. presented a total of 23 data sets for mercury capture by Norit’s DARCO-HG raw AC in a range of sorbent loadings of 0.01−0.062 g of AC/m3 of flue gas and a temperature range of 118−149 °C.5 The mercury removal efficiencies were reported on the basis of the measurements obtained under a steady-state condition during either 1 or 2 h. Because our fixed-bed experiments were carried out at 140 °C for the determination of adsorption parameters in our previous study, we have selected the 12 data sets obtained from the paper in a temperature range of 131−149 °C. Scala also summarized pilot-scale mercury removal results using FFs reported in the literature, as shown with the shaded areas in Figure 8. The 12 data points selected from the paper by Flora et al. between 131 and 149 °C showed a little higher mercury removal than the data summarized between 120 and 140 °C by Scala. However, these two sets of literature data were generally in good agreement. The adsorption parameters used in the paper by Flora et al. were determined by assuming very fast (instantaneous) adsorption kinetics (i.e., an adsorption isotherm was used). Because of this difference between our and their models, their model parameters could not be applied to our model to generate model predictions. However, because the Langmuir adsorption kinetics was also used in the paper by Scala, his model predictions were reproduced using the parameters for the HGR sorbent (sulfur-doped AC) used in our model.14 DARCO G60 (raw AC) was also used in the paper by Scala. However, its results were not extensively reported in the paper and, thus, were not used for comparisons. The model predictions made at 140 °C for 1 h using our own model were between those made for the HGR sorbent at 120 and 150 °C for 1 h, indicating reasonably good agreement. Overall, all model predictions were in good agreement with the pilot-scale experimental data presented in the paper by Scala, but somewhat underestimated the data in the paper by Flora et al. Throughout a comparison of the model predictions and pilot-scale data, in general, the following discrepancies were noted: (1) Our model presented in this study only describes HgCl2 adsorption. The pilot-scale data includes both Hg0 and HgCl2 species, but its speciation data was not clearly shown in the two previous papers. In addition, the models used in the two previous studies did not clearly describe the adsorption expressions for mercury species (i.e., Hg0 and/or HgCl2). (2)

There must be mercury adsorption onto the unburned carbon content of fly ash, but none of the models have such a description. (3) There must also be Hg0 capture as a result of in situ HCl adsorption onto the carbon surface (e.g., raw AC and unburned carbon in fly ash), which also leads to Hg0 vapor removal. However, these have not been taken into account in all of these models. Therefore, a more sophisticated model needs to be developed for more accurate model predictions by taking into account (1) the chemisorption of Hg0 vapor, such as Hg0 oxidation over CuCl2, followed by physical adsorption or in situ HCl adsorption onto the carbon surface followed by Hg0 adsorption and (2) the physical adsorption of oxidized mercury onto unburned carbon in fly ash.

4. CONCLUSION A modeling study for HgCl2 adsorption by sorbent injection by raw AC and 4 and 10% CuCl2−AC was conducted on the basis of the equilibrium and kinetic parameters determined from our previous study. Overall, the simulation results indicate that the filter cake can effectively capture HgCl2 in a FF by injecting the three sorbents. A baseline study of HgCl2 capture by raw AC (Norit DARCO-HG) injection in a FF system shows that an average 90% mercury removal efficiency can be achieved with a sorbent loading of 0.2 g/m3, a sorbent particle size of 20 μm, and a filtration time of 90 min. Our operation case study suggests that sorbent injection at a high loading for a short duration followed by the discontinuation of the injection until a cleaning cycle time can significantly save a total amount of a sorbent (e.g., >70%) to be injected and increase the sorbent utilization while achieving the same average mercury removal efficiency. Although the adsorption kinetic constant was found to slightly increase with CuCl2 loading, HgCl2 adsorption in the filter cake was predicted to be dominated by the adsorption capacity difference. The model predictions compared to pilotscale data suggest that the model needs to be further refined for more accurate predictions of field data.

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AUTHOR INFORMATION

Corresponding Author

*Telephone: 1-513-556-0018. Fax: 1-413-556-0018. E-mail: [email protected]. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS This study was funded by the National Science Foundation (NSF), CAREER Grant 1151017. The authors appreciate the financial support of the NSF. The authors also appreciate a Norit DARCO-HG AC sample provided by Norit Americas, Inc.

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7662

NOMENCLATURE C = HgCl2 concentration in the gas phase of the internal pore (g of HgCl2/m3) CB = HgCl2 concentration in the bulk gas phase of the filter cake (g of HgCl2/m3) CinB = HgCl2 concentration in the gas phase entering the filter cake (g of HgCl2/m3) 3 Cout B = HgCl2 concentration at the FF outlet (g of HgCl2/m ) 2 De = effective pore diffusion coefficient (m /s) Dp = sorbent particle diameter (m) K = equilibrium constant, K = k1/k2 (m3/g of HgCl2) dx.doi.org/10.1021/ef4017625 | Energy Fuels 2013, 27, 7654−7663

Energy & Fuels

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k1 = adsorption rate constant (m3 g−1 s−1) k2 = desorption rate constant (s−1) kg = gas-phase mass-transfer coefficient (m/s) L = filter cake thickness (m) LF = final cake thickness formed at a final filtration time (tF) mAC = sorbent injection loading in the flue gas (g/m3) mash = fly ash concentration in the flue gas (g/m3) q = HgCl2 adsorbing on the sorbent (g of HgCl2/g of sorbent) qavg = average HgCl2 uptake (g of HgCl2/g of sorbent) qmax = maximum adsorption capacity (g of HgCl2/g of sorbent) Rp = radius of the particle (m) r = sorbent radial distance (m) rA = HgCl2 net adsorption rate (g of HgCl2 adsorbed per m3 of sorbent per s) rA,obs = observed adsorption rate (g of HgCl2 adsorbed per m3 of sorbent per s) t = filtration time (s) tF = final filtration time (s) u = superficial gas velocity (m/s) z = axial filter cake thickness from the FF surface (m) ∗ = equilibrium state

(U.S. EPA): Washington, D.C., 2002; Vol. EPA/452/B-02-001, Section 6: Particulate Matter Controls, Chapter 1: Baghouses and Filters. (8) Jones, A. P.; Hoffmann, J. W.; Smith, D. N.; Feeley, T. J.; Murphy, J. T. DOE/NETL’s Phase II Mercury Control Technology Field Testing Program: Preliminary economic analysis of activated carbon injection. Environ. Sci. Technol. 2007, 41 (4), 1365−1371. (9) Srivastava, R. K.; Hutson, N.; Martin, B.; Princiotta, F.; Staudt, J. Control of mercury emissions from coal-fired in electric utility boilers. Environ. Sci. Technol. 2006, 40 (5), 1385−1393. (10) Zhao, B.; Zhang, Z.; Jin, J.; Pan, W.-P. Simulation of mercury capture by sorbent injection using a simplified model. J. Hazard. Mater. 2009, 170 (2−3), 1179−1185. (11) Skodras, G.; Diamantopoulou, I.; Pantoleontos, G.; Sakellaropoulos, G. P. Kinetic studies of elemental mercury adsorption in activated carbon fixed bed reactor. J. Hazard. Mater. 2008, 158 (1), 1−13. (12) Karatza, D.; Lancia, A.; Musmarra, D.; Pepe, F.; Volpicelli, G. Kinetics of adsorption of mercuric chloride vapors on sulfur impregnated activated carbon. Combust. Sci. Technol. 1996, 112 (1), 163−174. (13) Scala, F. Simulation of mercury capture by activated carbon injection in incinerator flue gas. 2. Fabric filter removal. Environ. Sci. Technol. 2001, 35 (21), 4373−4378. (14) Scala, F. Modeling mercury capture in coal-fired power plant flue gas. Ind. Eng. Chem. Res. 2004, 43 (10), 2575−2589. (15) Lee, S.-S.; Lee, J.-Y.; Keener, T. C. Mercury oxidation and adsorption characteristics of chemically promoted activated carbon sorbents. Fuel Process. Technol. 2009, 90 (10), 1314−1318. (16) Li, X.; Liu, Z.; Lee, J.-Y. Adsorption kinetic and equilibrium study for removal of mercuric chloride by CuCl2-impregnated activated carbon sorbent. J. Hazard. Mater. 2013, 252−253 (0), 419−427. (17) COMSOL, Inc. COMSOL Multiphysics, Version 4.3a; COMSOL, Inc.: Burlington, MA, 2012. (18) McKenna, J. D.; Nunn, A. B.; Furlong, D. A. In Air Pollution Engineering Manual, 2nd ed.; Davis, W. T., Ed.; Air and Waste Management Association: New York, 2000; pp 101−103. (19) Davis, W. R. Air Pollution Engineering Manual, 2nd ed.; John Wiley and Sons, Inc.: New York, 2000. (20) Donovan, R. P. Fabric Filtration for Combustion Sources; Marcel Dekker, Inc.: New York, 1985. (21) Li, X.; Lee, J.-Y.; Heald, S. XAFS characterization of mercury captured on cupric chloride-impregnated sorbents. Fuel 2012, 93, 618−624.

Greek Letters

εb = bed porosity εp = particle porosity ρp = sorbent particle density (g/m3) τ = dimensionless time (see eq 8) φ = dimensionless sorbent uptake (see eq 8) ξ = dimensionless radial distance (see eq 8) λ = dimensionless axial distance (see eq 8) γ = dimensionless pore concentration (see eq 8) γB = dimensionless bulk concentration (see eq 8) γout B = dimensionless outlet bulk concentration (see eq 8)

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REFERENCES

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dx.doi.org/10.1021/ef4017625 | Energy Fuels 2013, 27, 7654−7663