Numerical investigation of oxy-fuel combustion in 700 °C-ultra-supercritical boiler

Fuel 207 (2017) 602–614

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Full Length Article

Numerical investigation of oxy-fuel combustion in 700 °C-ultrasupercritical boiler Xueli Ge, Jiancong Dong, Haojie Fan, Zhongxiao Zhang ⇑, Xianyao Shang, Xinglei Hu, Jian Zhang School of Mechanical Engineering, Shanghai Jiao Tong University, 800 Dongchuan RD, Minhang District, Shanghai 200240, China

h i g h l i g h t s

g r a p h i c a l a b s t r a c t


New correlations for WSGGM are

proposed under oxy-fuel combustion.
Improved models are employed to

simulate oxy-fuel combustion in AUSC boiler.
The heat transfer in the furnace decreases by 4–11% under oxy-fuel combustion.
Combination of AUSC and oxy-fuel helps better arrange the heatexchanger surfaces.

a r t i c l e

i n f o

Article history: Received 9 March 2017 Received in revised form 1 May 2017 Accepted 27 June 2017

Keywords: Oxy-fuel combustion Advanced ultra-supercritical boiler Modelling Computational fluid dynamics

a b s t r a c t The combination of the advanced ultra-supercritical (AUSC) and oxy-fuel combustion technologies is considered the most feasible and promising choice for carbon capture and storage. The high efficiency of AUSC boilers compensates for the cost incurred in controlling CO2 emissions. In this study, a new improved model was proposed for gas-radiative properties as a polynomial function of the continuous molar ratio of H2O to CO2 and temperature, which was compiled as user-defined functions in C language. The modified radiative model was employed in the computational fluid dynamics analysis to simulate a tangentially fired 700 °C-USC boiler with a capacity of 300 MW under the oxy-fuel condition. The results show that the improved model has a good accuracy for gas-radiative properties compared to other classical models. In the oxy-fuel cases, the heat transfer obtained in the simulation is approximately 4–11% lower than that in the thermodynamic calculation because of the char gasification. The radiative capability of the flue gases in the oxy-fuel case is much greater than that in the air–fuel case with the increase in the molar fraction of oxygen. The heat transfer via the water wall is approximately 4–8% higher than that via air, when the molar fraction of oxygen increases from 26% to 29%. The increase in the heat absorption helps in better arrangement of the heat-exchanger surfaces in the furnace of the AUSC boiler. Ó 2017 Published by Elsevier Ltd.

1. Introduction The efficiency of power plants with a CO2 capture system decreases by 6–13%, and the cost incurred in capturing CO2 will

⇑ Corresponding author. E-mail address: [email protected] (Z. Zhang). http://dx.doi.org/10.1016/j.fuel.2017.06.119 0016-2361/Ó 2017 Published by Elsevier Ltd.

lead to an increase in the cost of electricity by approximately 20–30% [1,2]. Addressing the growing demand for electricity and controlling environmental pollution is a challenge. Nevertheless, a balance can be achieved by improving the efficiency of boilers and controlling the emission of CO2. Evidently, higher efficiency implies a lower emission of pollutants. In recent years, the oxyfuel combustion technology is regarded one of the most feasible

X. Ge et al. / Fuel 207 (2017) 602–614

technologies to capture and store CO2 emitted from fossil-fueloperated power plants, wherein the total cost of CO2 capture is approximately 60% compared to that in the amine case [2]. By analysing the energetic and exergetic comparison of oxy-fuel combustion, calcium looping and amine scrubbing, oxy-fuel combustion has competitive advantages in retrofitting for existing power plants [3,4]. In particular, in China and India, coal would be the dominant source of energy in the future owing to abundant reserves and low cost. Previous studies have shown that the gas shift from N2 to CO2 will lead to significant changes in the ignition location, heat transfer, and mechanism of reactions [5,6]. Recently, numerous research and developmental studies have been conducted and continuous progress has been made in the oxy-fuel-combustion technology [7–10]. However, the rising cost due to CO2 emission limits the development and commercial applications of oxy-fuel technologies. Increasing the operating temperature and pressure can clearly improve the efficiency of boilers [11,12]. More than two hundred 600 °C-ultra-supercritical (USC) boilers have been commercially built in China, demonstrating excellent availability and high efficiency [13]. For example, the unit thermal efficiency of the 1000 MW-Yuhuan-USC power unit is approximately 45.5%, and the boiler efficiency is approximately 93.88% [14]. The standard coal consumption of this unit is reduced to approximately 270.6 g/(kw h). The 600 °C-USC technology allows coal-fired power plants to generate more electricity and emit less pollutants compared to previous coal-fired plants. With the advancement in design, manufacture, and operating experience, the current objective is to investigate and develop advanced USC power plants with steam temperatures above 700 °C, for example, the European AD700 programme, the American A-USC (760 °C) programme, and the Japanese A-USC [15,16]. The CO2 emission from an A-USC power plant decreases by approximately 22% compared to conventional pulverised coal boilers [22]. However, substantial differences exist between USC and traditional coal-fired power plants, which are presented both in the fireside and steam-side scales, including the materials, characteristics of steam, endothermic ratio, and location of heatdelivery surfaces. The accuracy in determining the amount of heat transfer is critical for the design and retrofit of oxy-fuel power plants, which strongly depends on the parameters of radiative gases. Several studies have been conducted to improve the sub-models including the radiative-properties model and chemical-reaction mechanisms under the oxy-fuel combustion. The weighted sum of grey gases model (WSGGM) proposed by Smith is widely employed in conventional air-combustion systems [17]; however, the technology has not been applied to oxy-fuel combustion. Various attempts have been made to obtain new data and correlations for a reasonable prediction of oxy-fuel combustion [17–23]. Yin et al. [18] proposed new parameters for the WSGGM with four grey gases by employing the exponential wide-band model for temperatures in the range of 500–3000 K and pressure path lengths in the range of 0.001–60 bar m. The condition was divided into ten representative selection criteria. Johansson et al. [23] proposed new correlations for the WSGGM with four grey gases based on the statistical narrow-band model by employing the EM2C database in the temperature range of 500–2500 K and pressure path length in the range of 0.01–60 bar m. Kangwanpongpan et al. [21] developed new correlations for the WSGGM by fitting the emittance charts calculated from the advanced HITEMP database, which accounted for molar ratios (MRs) between 0.125 and 4, pressure path lengths from 0.001 to 60 bar m, and a temperature range of 400–2500 K. These models are not suitable for the simulation, because of the discontinuity of the MRs and absorption coefficients.

603

Computational fluid dynamics (CFD) modelling is used as an effective and vital tool to analyse furnace combustion [24–29]. Many studies have employed the air–fuel WSGGM to investigate the radiation heat transfer using the FLUENT code [30–34]. However, oxy-fuel combustion is fundamentally different from air combustion, particularly in terms of the radiation heat transfer in the furnace because of the higher concentration and different MRs of H2O to CO2. Because of the different compositions and quantity of flue gases between air and oxy-fuel combustions, the characteristics of the heat transfer are significantly different. Therefore, models that accurately describe the radiation properties are essential to analyse the heat transfer in the furnace. However, the existing models are only suitable for the classical conditions (e.g. MR = 0.125 or 0.25), which are discontinuous for MRs. Generally, the absorption coefficients are assumed constant or only as a function of variable MRs. In the real furnace, the MR is continuous and not discrete. Consequently, the discrete MR and absorption coefficients cannot be incorporated in the CFD, unless the combustion conditions are divided into classical conditions [18]. Therefore, it is very essential to build a new model, which can propose the gas radiation properties by continuous MRs, especially for the value of emittances and absorption coefficient. This study provides a numerical simulation with a new set of correlations for the WSGGM to investigate the feasibility and advantage of combining the USC and oxy-fuel technologies in the furnace. The new correlations are determined by fitting the total emittances of the mixture gases including CO2 and O2. To improve the accuracy of gas-radiative heat transfer, the absorption coefficients are considered as a function of the continuous MR and temperature. The new model is then validated by comparing the emittances with other classical models and demonstrating the CFD modelling of a tangentially fired AUSC boiler of capacity 300 MW. The difference and trends of heat transfer between the oxy-fuel and air–fuel combustions are proposed in this study. This study provides a quantitative assessment of the advanced USC boiler for the design of oxy-fuel combustion systems.

2. Boiler description and operating conditions 2.1. Boiler description and mesh generation A 300-MW tangentially fired USC boiler is modelled, as shown in Fig. 1(a), which is a representative pulverised-coal boiler. The geometrical parameters of the boiler are 12,350, 14,022, and 50,868 mm in terms of depth, width, and height respectively. Twenty burner nozzles are arranged at each corner, including six primary air (PA) nozzles, nine secondary air (SA) nozzles, two over fir air (OFA) nozzles, and three separated over firing air (SOFA) nozzles. The entire furnace is divided into two zones by the OFA, main combustion, and burn-out zones, wherein the excess air ratio changes from approximately 0.85 to 1.25 in the conventional air– fuel combustion. The SA is injected into the furnace with a contrary tangent of 17° in the direction of the PA to improve the mixing of the pulverised coal and air. The OFA is injected into the furnace with the same tangent of 20° to reduce the swirl intensity of the combustion gases. In the upper furnace, four groups of division platen superheaters and twenty rear panels with an equal interval are installed. This boiler selected in this paper is a typical seriation of industrial steam boilers of China. The research zone covers the areas from the ash hopper to the vertical flue pass, which is generated by a partition meshing method with a hexahedral-structured mesh. A grid independence test are proposed to clean up the minimum cell to ensure the accuracy of the simulation results. The results of average cross-section temperature distribution along the flue gas direction show the

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X. Ge et al. / Fuel 207 (2017) 602–614

Fig. 1. Schematic of cross section of furnace and mesh. (a) Furnace geometry, (b) y = 7.01 m, (c) z = 6.97 m.

maximum deviation is about 3.27% in the 971,338 cells and 6.52% in the 323,780 cells compared to the results of the 462,542 cells. Finally, the optimised mesh containing 462,542 cells is selected to balance the accuracy and computational cost. Fig. 1(b) and (c) shows the cross section of the mesh system.

2.2. Coal properties and case study The coal employed in this study is bituminous coal with a size in the range 60–120 mm. The particle-size distribution is assumed to follow the Rosin–Rammler algorithm with a mean size of 95 mm and a spread parameter of 3.5. Table 1 presents the proximate and ultimate analysis of the coal particles. The coal ignites easily owing to the high volatility and low moisture. To compare the oxy-fuel combustion with the conventional air– fuel combustion, the inlet gases of each nozzle for the oxy-fuel combustion should be similar to those of the air–fuel combustion. To investigate the performance of the oxy-fuel combustion, the air–fuel combustion is selected as a reference, and three other cases are employed: Oxy21 (21 vol.% O2), Oxy26 (26 vol.% O2), and Oxy29 (29 vol.% O2). Table 2 lists the detailed parameters of the flow. The numerical investigation helps provide quantitative analysis of the heat transfer and temperature distributions with

Table 1 Proximate analysis, ultimate analysis and calorific value of coal particle. Parameter

Units

Value

Proximate analysis (as received) Fixed carbon Volatiles Ash Moisture

% % % %

42.87 16.57 34.56 6.0

% % % % %

50.25 3.08 4.25 0.93 0.93

kJ kg1

19,230

Ultimate analysis (as received) C H O N S Calorific value (as received) Low heating value

the increase in O2. To obtain a high flue temperature to achieve a steam temperature of 700 °C, the PA temperature should be much higher than that in the conventional subcritical boiler (approximately 343 K). The air–fuel case corresponds to the working power plant. The boundary conditions of the inlet gases for the three oxyfuel cases are confirmed based on the momentum of the air–fuel case.

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X. Ge et al. / Fuel 207 (2017) 602–614 Table 2 Boundary conditions for full-scale boiler simulations.

Table 3 Coefficients cki,j of four grey gases for WSGGM.

Cases

PA flow ratio (%) PA temperature (K) PA velocity (m/s) SA flow ratio (%) SA temperature (K) SA velocity (m/s) Coal flow rate (ton/h) Excess oxygen ratio

Air–fuel

Oxy21

Oxy26

Oxy29

18.4 602 23 76.6 593 48 165.15 1.25

18.4 602 23 76.6 593 48 165.15 1.25

22.8 602 23 72.2 593 48 165.15 1.25

25.3 602 23 69.2 593 48 165.15 1.25

3. Numerical modelling The commercial CFD code Fluent (version 14.5) is used to simulate the pulverised-coal combustion process in the furnace, including the turbulent flow, volatile combustion, char oxidation and gasification, and heater transfer. The realisable k-epsilon model is employed to simulate the turbulent flow with a correction of buoyancy effect. The trajectory of the coal particles is simulated using a discrete phase model considering the effects of the drag and gravity forces. The char combustion is simulated using the finite-rate/eddy-dissipation for the turbulence-chemistry interaction. The discrete ordinates (DO) model is used for radiation heat transfer. Because of the change in atmosphere from N2 to CO2, the following three models should be modified for the oxy-fuel case: radiative model, char-oxidation model, and homogeneousvolatile-oxidation model. 3.1. Radiative model In the furnace, the radiative model is the dominant model of heat transfer, including particle and gas radiative behaviours. Because of the gas shift in the oxy-fuel case, gas radiative parameters are very important to accurately evaluate the heat transfer. Hottel and Sarofilm proposed a WSGGM [35], which can be applied in the CFD calculations. It is assumed that the radiation can be obtained by a set of Ng grey gases and another clear gas, wherein each gas has an individual absorption coefficient. The total emittances are functions of the temperature, pressure path length, and absorption coefficient.

e¼

Ng X wi ð1  expðki LPðY CO2 þ Y H2 O ÞÞÞ

ð1Þ

i

j

c1i,j

c2i,j

c3i,j

1 1 1 1 2 2 2 2 3 3 3 3 4 4 4 4

1 2 3 4 1 2 3 4 1 2 3 4 1 2 3 4

0.070113 0.08225 0.03914 0.00661 0.45044 0.52054 0.259516 0.04855 0.059645 1.099669 0.95508 0.221254 0.419802 0.49688 0.656425 0.16609

0.254835 0.39399 0.221961 0.04287 0.49339 2.391296 2.02922 0.489448 0.01961 0.94546 1.064002 0.28039 0.258166 1.05184 0.743259 0.16619

0.05795 0.142473 0.1044 0.023894 0.124878 0.8929 0.77061 0.18627 0.040412 0.296147 0.33584 0.08682 0.10734 0.454277 0.33037 0.075552

ted the correlations for the emittances when the temperature was approximately 2000 K, compared to 1200 K in the Refs. [18,21,22].

wi ¼

ð4Þ

j¼1

where cij denotes the polynomial coefficients. The polynomial coefficients cij, number of polynomial coefficients N, number of grey gases Ng, and reference temperature Tref can be obtained by fitting the absorption coefficients and weighting factors to the total emittances calculated using line-by-line (LBL) or other models from the HITEMP 2010 database [36]. Generally, ri is assumed constant and cij depends only on the MR as the independent variable [21,23]. Four absorption coefficients exist for each typical MR. Yin et al. developed new WSGGM parameters and applied them to the representative oxy-fuel conditions with a constant MR, wherein the condition was divided into ten selection cases [18]. However, the MR varies successively from the hopper to the furnace exit in real combustion. The method proposed by Yin can be conveniently implemented in the oxy-fuel combustion simulation by compromising the prediction accuracy. The results show that the deviation in the model proposed by Yin is the highest among the six prediction models [21]. To improve the accuracy and reduce the deviation and computational cost, a set of ri and cij is determined not only by using the MR of the second order but also by using the temperature-dependent polynomial function of the third order, which can be expressed as follows.

i¼0

Here, wi is the weighting factor, and the summation of wi is equal to unity. Thus, the weighting factor of the window gas (i = 0) can be obtained as follows.

ki ¼

Ng X w0 ¼ 1  wi

bij ¼

ð2Þ

N X cij ðT=T ref Þj1 ;

N¼4 X bij ðT=T ref Þj1

ð5Þ

j¼1 3 X bk;ij MRk1

ð6Þ

k¼1

i¼1

The weighting factor wi is calculated by employing temperature polynomials T with the associated polynomial coefficient bij as follows:

wi ¼

N X wi ¼ bij T j1 ;

cij ¼

ð3Þ

N¼4 X cij T j1

ð7Þ

j¼1 3 X ck;ij MRk1

ð8Þ

k¼1

j¼1

where bij depends on the temperature or molar ratio of H2O/CO2. The prediction accuracy can be improved by normalising the temperature with the reference temperature Tref. Moreover, the normalised temperature helps simplify the optimisation procedure [21]. Tannin et al. concluded that the reference temperature best fit-

In Eqs. (6) and (8), the MR is varying from 0.125 to 4.0, including in the almost wet and dry-flue-gas-recycle (FGR) cases. The nonlinear multiple-regression analysis and Levenberg–Marquardt algorithm were employed to obtain the new correlations in this study. The MR and temperature-dependent polynomial coefficients cki,j and bki,j were determined by fitting the total emittances,

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X. Ge et al. / Fuel 207 (2017) 602–614

Table 4 Coefficients bki,j of four grey gases for WSGGM.

Table 5 Mechanisms of Char surface reactions.

i

j

b1i,j

b2i,j

b3i,j

1 1 1 1 2 2 2 2 3 3 3 3 4 4 4 4

1 2 3 4 1 2 3 4 1 2 3 4 1 2 3 4

105.3082 0.017262 0.01525 0.003911 3.386233 1.693028 1.19358 0.266963 0.06371 1.173845 0.9532 0.232854 0.007763 0.028807 0.02902 0.007428

39.1954 0.3669 0.280155 0.06612 0.83532 4.250212 3.33618 0.799926 1.191368 4.86797 4.197609 1.05866 0.013514 0.02448 0.026231 0.00774

6.043568 0.23771 0.17932 0.041997 0.678994 3.00472 2.296622 0.54187 0.54334 2.702553 2.27965 0.568522 0.00993 0.030806 0.02717 0.00698

which covered the MR from 0.125 to 4.0 and pressure-path length range of 0.01–60 m in the temperatures between 500 and 2600 K. The fitting procedure of the correlations with total emittances is summarised below: (1) the total emittances were calculated from the LBL integration of the HITEMP 2010 database at MRs = 0.125, 0.25, 0.75, 1.0, 1.25, 1.5, 2.0, and 4.0. Twenty-three temperatures were selected from 400 K to 2600 K with a step of 100 K. The path length was also divided into ten values: 0.01, 0.1, 1.0, 5.0, 10.0, 20.0, 30.0, 40.0, 50.0, and 60.0. (2) the initial values of the absorption coefficients ki were assumed as constant. For each emittance, ki and wi can be obtained using the Levenberg–Marquardt algorithm as expressed in Eqs. (5) and (7). (3) cki,j and bki,j were also determined using a non-linear multiple-regression method expressed in Eqs. (6) and (8). The correlations were determined using the non-linear multiple-regression analysis in the codes of FORTRAN. Ultimately, the number of grey gases Ng = 4 and reference temperature Tref = 1200 K were found to be the best choice to balance the computational cost and accuracy. Tables 3 and 4 list the new correlations. The new WSGGM is compiled as user-defined functions (UDF) in the C language to replace the original Smith’s WSGGM in Fluent. 3.2. Char-oxidation model The particle temperature has a single peak and a sharp drop in the oxy-fuel combustion identified via experimental observation and pyrometer-measured data [9,37,38]. The gasification of char with the CO2 or H2O is essential to be included in the combustion-process simulation. In addition, different adiabatic flame temperatures correspond to different combustion mechanisms for the oxy-fuel combustion. The deviation in the adiabatic flame temperature is approximately in the range 200–500 K for different mechanisms [36]. The global-combustion mechanism can be defined as ‘1-step’, ‘2-step’, and ‘4-step’ [9,37–42]. In 1step, CO2 and H2O are the only products of the reaction; CO, CO2, and H2O are the products for 2-step; and H2, CO, CO2, and H2O are the products for 4-step. In this study, the ‘4-step’ global-

Number

Reaction

A kg m2 s1 Pa1

E J/kmol

0 1 2

C + 0.5O2 ? CO C + H2O ? H2 + CO C + CO2 ? 2CO

0.005 0.002 0.00635

7.4  107 1.47  108 1.6  108

combustion mechanism is employed to investigate the combustion process in the furnace, displayed in Table 5. 3.3. Homogeneous-volatile-oxidation model Massive amounts of hydrocarbons derived from the volatile gases arising during the pyrolysis of the pulverised coal burn homogeneously in the gas phase. Because of the shift of N2 to CO2 and H2O, the concentration of CO will increase dramatically because of the reductions in CO2 and H2O. For the oxy-fuel combustion, the parameters of the homogeneous-reaction mechanisms should also be modified to apply to the new model. However, the primary products generated from the char-surface reactions burn further in the bulk gas, as given in Table 6: (0), (2), and (3) for oxy-fuel; (0), (1), and (2) for air-fired. 4. Results and discussion 4.1. Radiative properties To evaluate the accuracy of the new correlations, the total calculated emittances are compared with the benchmark LBL based on HITEMP 2010 database and other classical models. The results calculated using the correlations proposed by Johansson et al. [23], Tanin et al. [21], Yin et al. [18] are denoted as Johansson, Tanin, and Yin, respectively, in Figs. 2(a) and (b) and 3 (a) and (b); the results of this study are denoted by ‘new correlation.’ Fig. 2(a) and (b) shows the total emittances calculated using the different correlations with the path length varying from 0.01 to 60 m for MRs of 0.125 and 1.0, respectively. With the increase in the path length, the total emittances continuously increase and gradually tend to one. The figures show that the correlations have the same tendency. For shorter path lengths (L < 5 m), the error in the total emittances calculated by other authors can be neglected. For path lengths greater than 5 m, the error will continuously increase. As presented, the new correlations of this study and those obtained by Tanin provide the best accuracy compared to the reference LBL-HITEMP 2010 solution for the entire path-length range, followed by the correlations proposed by Johansson and Yin. The correlations of this study result in the higher value to the reference value, with the highest deviation of approximately 9.7% for MR = 0.125 and 5.3% for MR = 1.0, respectively. The average deviations are 6.24% and 3.04% for MR = 0.125 and MR = 1.0, respectively. Moreover, the average deviations are 12.2%, 19.1%, and 13.5% for the other three correlations when MR = 1. Fig. 3(a) and (b) illustrate the total emittances calculated using different correlations with respect to temperature varying from

Table 6 Homogeneous-reaction mechanisms of gas-phase reactions [40,43]. No.

Reaction

Rate equation (kmol m3 s1)

A kg m2 s1 Pa1

E J/kmol

0 1 2 3

VM + O2 ? CO + H2O + N2 + SO2 CO + O2 ? CO2 H2 + O2 ? H2O CO + O2 ? CO2

d[CO]/dt = ATbexp(-E/RT)[VM]0.2[O2]1.3 d[CO]/dt = ATbexp(-E/RT)[CO][O2]0.25[H2O]0.5 d[H2]/dt = ATbexp(-E/RT)[H2][O2] d[CO]/dt = ATbexp(-E/RT)[CO][O2]0.25[H2O]0.5

2.12  1011 2.24  108 9.87  108 2.24  106

2.03  108 1.7  108 3.1  107 4.2  107

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X. Ge et al. / Fuel 207 (2017) 602–614 1.0

0.9

Yin Tanin Johansson New correlation LBL

P=1.0bar,MR=0.125 0.8

0.8 0.7

Emittance

Emittance

0.6

0.4

Yin Tanin Johansson New correlation LBL

0.2

P=1.0bar,MR=0.125 0.0

0.6 0.5 0.4 0.3 0.2

0

10

20

30

40

50

60

0

500

1000

1500

2000

2500

3000

2500

3000

Temperature (K)

Path length (m)

(a)

(a) 0.9

1.0 0.8

0.8

Emittance

0.7

Emittance

0.6

0.4

Yin Tanin Johansson New correlation LBL

0.2

P=1.0bar,MR=1 0.0

0.6

0.5

Yin Tanin Johansson New correlation LBL

0.4

0.3

0

500

1000

P=1.0bar,MR=1 1500

2000

Temperature (K) 0

10

20

30

40

50

60

(b)

Path length (m)

(b)

Fig. 3. Total emittances with respect to temperature for different correlations of WSGGM at L = 10 m for (a) MR = 0.125 and (b) MR = 1.

Fig. 2. Total emittances with respect to path lengths for different correlations of WSGGM at T = 1500 K for (a) MR = 0.125 and (b) MR = 1.

Absorpation coefficient,k (1/(bar .m))

500 to 2600 K at L = 10 m for MR = 0.125 and 1.0, respectively. The figures shows that the total emittances continuously decline with the increase in temperature. Generally, the new correlations of this study and those obtained by Tanin provide the best accuracy compared to the benchmark LBL-HITEMP 2010 solution for all the temperatures, where average deviations of approximately 1.41% and 1.18%, respectively, for MR = 1.0 are observed. Furthermore, the average deviation calculated by Johansson and Yin’s correlations is 13.57% and 15.33%, respectively. The same change tendency of the total emittances is observed when MR = 0.125. The average deviations for MR = 0.125 are 2.56%, 3.7%, 14.41%, and 18.46% for the correlations of this study, Tanin, Johansson, and Yin, respectively. The results show that the new modified correlations for WSGGM yield a good accuracy of the total emittances for the mixed combustion gases. The temperature, path length and MR of demonstration are all distributed within the scope of the valid range verified in this section. In Fig. 4, the new polynomial coefficients expressed in equations of (6) and (7) are in good agreement with the absorption coefficients obtained from the LBL calculations adopting HITEMP 2010 database for MRs between 0.125 and 4.0. It is practical to adopt the MR and temperature as the variables for absorption coefficients.

1000

100

10

1

0.1

κ1 κ2 κ3 κ4

0.01

1E-3

0.0

0.5

1.0

1.5

2.0

2.5

k1-new k2-new k3-new k4-new 3.0

3.5

4.0

4.5

MR Fig. 4. Absorption coefficients of each grey gas obtained from the new correlation of WSGGM in Eq. (6) (lines) compared to the absorption coefficients from the Ref. [21] (symbols).

Fig. 5 shows the detailed comparison of the emittances between the air–fuel and oxy-fuel calculated using the correlations proposed in this paper and by Smith et al. There is a considerable

608

X. Ge et al. / Fuel 207 (2017) 602–614 0.60 0.55 0.50

Emittance

0.45 0.40 0.35 0.30 0.25

Air-fuel-Smith Air-fuel-new Oxy29-Smith Oxy29-new

0.20 0.15 0.10 -10

-5

0

5

10

15

20

25

Height (m) Fig. 5. Emittances calculated from Smith and new WSGGM.

30

decrease in emittances from the ash hopper to the platen superheaters in the upper furnace. For the air–fuel case, the two models are in good agreement along the height of the furnace with an average deviation of 4.7%. The total emittances change abruptly with the increase in height. The change in the curve can be clearly observed at heights of approximately 4.5 m and 18.0 m, i.e. between the primary combustion (4.289–11.469 m) and SOFA zones (18.429–20.269 m). The fluctuations in the emittances curves are mainly due to the inlet of the reactants and combustion. Although a similar trend appears in the case Oxy29 along the height, there is a completely different relationship between the new and Smith models. In contrast to the air–fuel case, the value calculated using the new correlations are higher than that of Smith with a maximum deviation of 20.4% at the location of the main combustion zone. Moreover, emittances vary moderately and have a completely different trend compared to the air–fuel cases in the whole furnace. The new emittances of the oxy-fuel cases are lower in the hopper and main combustion zone, but higher than those in the air–fuel case over the location of SOFA. The above changes in

(a) Air-fuel

(b) Oxy21

(c) Oxy26

(d) Oxy29

Fig. 6. Distributions of flue-gas temperature in cross section for different cases (height = 6.97 m).

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X. Ge et al. / Fuel 207 (2017) 602–614 1700 1800

1600 1500

Air-fuel Oxy21 Oxy26 Oxy29

1600

Temperature (K)

Temperature (K)

1400 1300 1200 1100

Air-fuel Oxy21 Oxy26 Oxy29

1000 900 800 0

1200

1000

800

10

20

30

40

-10

30

(b) centreline of furnace

1800

400

Heat Transfer (MW)

1600 1500 1400

Air-fuel Oxy21 Oxy26 Oxy29

1300 1200

0

20

(a) front wall 420

-5

10

Height (m)

1900

1100 -10

0

Height (m)

1700

Temperature (K)

1400

5

10

15

20

25

40

50

Simulation Thermodynamic calculation

380 360 340 320 300

30

case

air

oxy21

oxy26

oxy29

Height (m)

(c) cross section

(d) heat transfer in furnace

Fig. 7. Temperature distributions at different positions along the height and heat transfer in furnace.

emittances may be because of the following two reasons: (1) the high molar fraction and substantial number of CO2 and H2O molecules and (2) the significant role of flue gas temperatures in influencing the radiative properties in the air–fuel cases. 4.2. Temperature distributions Fig. 6 shows the temperature distributions in the cross section (at height = 6.97 m) for all the cases. This location is the fourth pulverised coal burner from the ash hopper. The mean temperature of this cross section is very close to the average temperature of flame in the main combustion zone. As given in Table 2, the inlet-gas temperature of the PA and SA are 602 K and 593 K, respectively. For Oxy21, the highest temperature declines to 1760 K, because N2 is replaced by CO2. With the increase in the molar fraction of oxygen, the highest temperature rises gradually. When the molar fraction reaches 29%, the highest temperature is approximately 2050 K, exceeding that in the air–fuel case (2020 K). The results have a good agreement with the results obtained by Al-Abbas [44]. This can be attributed to the changes in the mass flow of the products and specific heat of the flue gases. The heat transfer and endothermic ratios in the water wall sharply improve with the increase in O2 molar fraction. The endothermic ratio may exceed 42.85% to meet the heat absorption requirement for the 700 °C-USC boiler through increasing the molar fraction of O2. Moreover, the diameter of the combustion area in the cases of Oxy26 and Oxy29 is larger

than that in the air–fuel case because the flow rate of the flue gases decreases considerably. Furthermore, the width of the oxy-fuel combustion zone generated by the fuel injected via the PA nozzles of each corner is significantly narrower than that in the air–fuel case. This phenomenon may be because of the ignition delay and increase in specific heat [45]. Fig. 7(a)–(c) shows the temperature profiles along the furnace centreline for all the cases, including front-wall temperature, centre temperature, and mass-averaged temperature of the cross sections. The value increases gradually from the hopper to the main combustion zone because of the fuel combustion, and subsequently, drops abruptly because of the inlet oxidant via the SOFA. As the flue gases rise up in the furnace, the heat-transfer process between the flue gases and water wall leads to the decline in the flue temperature. The gas temperature is a balance between the heat released during combustion and heat absorption of the water wall. The tendency of each case is similar along the height of the furnace. Compared to the benchmark air–fuel case, the temperature of Oxy21 is lower by approximately 240 K in the main combustion zone. The differences between the oxy-fuel and air–fuel cases are approximately 0.72% and 0.63% for Oxy26 and Oxy29, respectively. The similar difference between the oxy-fuel and air–fuel cases is also observed in the front-wall temperature and centre temperature. These results are consistent with those of another study [44]. Fig. 7 shows that the temperature gradually increases with the increase in oxygen, which has a good agreement with the results presented in Fig. 6.

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(a) Air-fuel

(b) Oxy21

(c) Oxy26

(d) Oxy29

Fig. 8. Char-mass fraction field of particles under air and oxy-fuel combustion cases.

For the operating conditions of the AUSC (36.48 MPa/705 °C/723 °C), the heat-absorption ratio of the steam/water at temperatures below 500 °C is approximately 57.15% compared to the entire heat-absorption process. Furthermore, the ratio would reduce to 48% for the double-reheat AUSC boiler. The ratio of the heat absorption through the water wall would gradually decline with the increases in temperature and pressure. Generally, it is approximately 1300 °C at the entrance of the convection-heatdelivery surface. The energy required to heat the steam to 700 °C is insufficient, which is less than 40% of the total heat absorption. Thus, it is necessary to obtain more energy in the furnace by improving the intensity of the heat transfer or moving tailheating surfaces into the furnace. However, the boiler tube is more prone to failure because of the steam-side oxidation and fireside corrosion mechanisms, as the steam temperature increases [46,47]. Fig. 7(d) shows the detailed comparison of the heat transfer between the simulation and design calculations simultaneously. The accuracy of numerical models is verified by the results

of simulation and thermodynamic calculation for air–fuel case. For the oxy-fuel cases, the simulation value is below the range 12–45 MW (4–11%), and the difference between them increases with the increase in the molar fraction of oxygen. The decline of the adiabatic flame temperature due to the char gasification may explain this deviation, which is neglected in the traditional thermodynamic calculation. Therefore, it is necessary to consider this deviation of heat transfer in the design procedure. With the increasing of oxygen concentration, the heat transfer of Oxy29 is approximately 4–8% higher than that of the air–fuel case, which can better satisfy the insufficient heat absorption of the heatdelivery surfaces by the water wall for the 700 °C-USC boiler. As mentioned, more heat exchange surfaces are needed to absorb sufficient energy to gain the high rated parameters of steam in the tail flue. However, the heating surfaces located in the tail flue is too compact that there is not enough space for the new surfaces. The unit heat absorption of the water wall can be enhanced by increasing the concentration of oxygen and the area of the water wall will

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(a) Air-fuel

(b) Oxy21

(c) Oxy26

(d) Oxy29

Fig. 9. Temperature distributions of particles under air and oxy-fuel combustion cases.

relatively decrease. The unnecessary surface of water wall can be replaced by the tail-heating surfaces. Therefore, the increase of heat absorption in the furnace will help improve the arrangement of the heat-exchanger surfaces, which will make the arrangement of heat-exchanger surfaces more reasonable. Fig. 8 shows the char-mass-fraction field of the particles under the oxy-fuel and air–fuel cases. The pulverised coal injected into the furnace from 1# corner is selected as the research subject. For all cases, it can be clearly observed that six layers of the pulverised coal entering into the furnace based on the tangential direction rotate upward. The pulverised coal particle burns out at the location of the arch nose in the air–fuel case. For Oxy21, the ignition delay can be easily observed in the results, and the burn

out position moves up to the outlet of the arch nose, thereby causing fouling and slagging of the superheating surfaces. The above simulation results correspond to the temperature profiles shown in Fig. 7(c). As shown in the Ref. [48], temperature and wall radiative fluxes could be replicated through an appropriate O2 fraction. Consequently, increased oxygen fraction evidently improves the ignition delay and decreases the elevation of the burnout. When the molar fraction of oxygen is 26%, Oxy26 has a same burn out position as that in the air–fuel case. It is generally accepted that radiative heat transfer is dominant in the furnace, and the high-temperature particles are the main source of the radiative heat transfer [49]. Therefore, it is necessary to obtain the temperature distribution of the particles in the

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(a) Air-fuel

(b) Oxy21

(c) Oxy26

(d) Oxy29

Fig. 10. Residence time of particles under air and oxy-fuel combustion cases.

furnace. Fig. 9 shows the temperature distribution of the particles for all the cases. When the molar fraction of oxygen is 21% or 26%, the particle temperature is clearly lower than that in the air–fuel case at the same position because of the char gasification and specific heat capacity of CO2. The simulation results explain why the heat transfer of the oxy-fuel cases is lower than that in the benchmark air–fuel case shown in Fig. 7(d). The Oxy29 case can better match the temperature distribution of the particles in the

air–fuel case. The ignition delay is also observed in the oxy-fuel cases via the particle-temperature distributions. Hence, only a slight difference exists in the heat transfer between the Oxy26 and air–fuel cases, which can be used for retrofit. Oxy29 is more suitable for AUSC because of the improved heat transfer. The residence time of particle is very important aspect for the heat transfer and burnout rate, which is an influential indicator of boiler efficiency. Fig. 10 displays the spatial distribution of

X. Ge et al. / Fuel 207 (2017) 602–614

particle residence time for all the examined cases. Clearly, the residence time increases for all oxy-fuel cases compared to the air– fuel case, because of the decrease of flue gas. From the ash hopper to the upper of the furnace, the residence time of particle under air–fuel is about 4.8 s, which raise to 5.175, 5.695 and 6.041 s for three oxy-fuel cases, respectively. Therefore, it is proved form another view that the oxy-fuel combustion can well make up the insufficient heat absorption by tail-heating surfaces. 5. Conclusions In this study, new correlations for the WSGGM are proposed as functions of the continuous MR and temperature. These parameters were fitted by employing the emittances calculated based on the timely HITEMP 2010 database. The improved model yields a better accuracy in terms of the total emittances for the mixed combustion gases with a highest deviation of approximately 9.7%. This model can be employed in the CFD code to investigate the real industrial oxy-fuel combustion process in the temperature range of 400–2600 K and pressure path length between 0.01 and 60 bar m. Several assumed cases have been investigated to evaluate the accuracy of the models simultaneously. The new WSGGM has been employed to CFD modelling of the combustion process for a 300-MW tangentially fired AUSC boiler by employing the UDF in the C language code. The temperature distributions of the flue gases and particle parameters are proposed to analyse the difference between the oxy-fuel and air–fuel conditions. For the same O2 molar fraction, temperature of oxy-fuel case is lower than air–fuel case. However, the radiative capability of the flue gases in the oxy-fuel cases is much greater than that of the air–fuel case as the concentration of oxygen increases. Moreover, the heat transfer in the oxy-fuel case in the main combustion zone is approximately 4–11% lower than that obtained via thermodynamic calculation. Oxy29 is more suitable for the design of 700 °C-USC boiler, when Oxy26 can be used for the existing unit retrofit. The increase in heat absorption in the furnace will help improve the arrangement of the heat-exchanger surfaces of the AUSC. The results show that the combination of the 700 °C-USC and oxy-fuel technologies will help achieve a reasonable arrangement of the heat-exchanger surfaces in the furnace. Acknowledgments This work was funded by the National Natural Science Foundation of China – China (Grant No. U1361201). References [1] Gambini M, Vellini M. CO2 emission abatement from fossil fuel power plants by exhaust gas treatment. J Eng Gas Turbines Power 2003;125:365–73. [2] Singh D, Croiset E, Douglas PL, Douglas MA. Techno-economic study of CO2 capture from an existing coal-fired power plant: MEA scrubbing vs. O2/CO2 recycle combustion. Energy Convers Manage 2003;44:3073–91. [3] Doukelis A, Vorrias I, Grammelis P, Kakaras E, Whitehouse M, Riley G. Partial O2-fired coal power plant with post-combustion CO2 capture: a retrofitting option for CO2 capture ready plants. Fuel 2009;88:2428–36. [4] Atsonios K, Panopoulos K, Grammelis P, Kakaras E. Exergetic comparison of CO2 capture techniques from solid fossil fuel power plants. Int J Greenhouse Gas Control 2016;45:106–17. [5] Khare S, Wall T, Farida A, Liu Y, Moghtaderi B, Gupta R. Factors influencing the ignition of flames from air-fired swirl pf burners retrofitted to oxy-fuel. Fuel 2008;87:1042–9. [6] Kakaras E, Koumanakos A, Doukelis A, Giannakopoulos D, Vorrias I. Oxyfuel boiler design in a lignite-fired power plant. Fuel 2007;86:2144–50. [7] Scheffknecht G, Al-Makhadmeh L, Schnell U, Maier J. Oxy-fuel coal combustion—A review of the current state-of-the-art. Int J Greenhouse Gas Control 2011;5:S16–35. [8] Wall T, Liu Y, Spero C, Elliott L, Khare S, Rathnam R, et al. An overview on oxyfuel coal combustion—state of the art research and technology development. Chem Eng Res Des 2009;87:1003–16.

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