Particle sticking behavior near the throat of a low-NOx axial-swirl coal burner

Applied Energy 88 (2011) 650–658

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Particle sticking behavior near the throat of a low-NOx axial-swirl coal burner Zhengqi Li ⇑, Lingyan Zeng, Guangbo Zhao, Shanping Shen, Fucheng Zhang School of Energy Science and Engineering, Harbin Institute of Technology, 92, West Dazhi Street, Harbin 150001, PR China

a r t i c l e

i n f o

Article history: Received 1 May 2010 Received in revised form 21 July 2010 Accepted 19 August 2010 Available online 16 September 2010 Keywords: Swirl burner Numerical simulation Particle sticking behavior

a b s t r a c t The results of numerical simulations of particle sticking behavior near the throat of a low-NOx axial-swirl burner in a 600-MWe bituminous coal burned boiler are presented. A comparison of simulation results with measurements using a probe with hot-film sensors shows that the numerical model offers a reasonable description. Calculated results of slagging show that slag build-up is substantial near the throat of the designed burner and that the sticking-particle ratio is as much as 33.2%. Because central and primary air streams remain unchanged, the mass flux of the inner secondary air is 3.25 kg s1 while that of the outer secondary air is 12.16 kg s1; however, the sticking-particle ratio can still be lowered to as little as 9.6%. By adjusting the outer secondary air blade angle to 15°, the sticking-particle ratio can be further lowered to 8.73%. Ó 2010 Elsevier Ltd. All rights reserved.

1. Introduction Slagging, a perennial problem associated with coal-fired boilers, refers to molten deposits within the furnace in areas directly exposed to flame radiation emanating from, for example, furnace walls and widely-spaced pendent super-heaters. During boiler operations, four types of problems associated with slag can arise [1]: (1) slag on the water-cooled tube wall of the boiler can reduce the radiation heat transfer from the furnace and the flue temperature in the super-heater region can increase. This results in an increase in the steam temperature in the super-heater that requires water to be sprayed into the super-heater; (2) slag on the wall can lead to increased corrosion; (3) slag can drop to the bottom of the furnace, causing damage to the tubes, leading to hopper blockages or difficulties in grinding the bottom ash; and (4) slag in the burner vent area can partially plug the throat, creating various problems such as a reduction in the combustion zone and damage to the bottom of the furnace from falling slag. Slagging characteristics are influenced by coal types (ash constituents, melting temperature, distribution of mineral matter, etc.), reaction atmosphere, particle temperature, surface temperature of heat exchanger, flow dynamics and so on. A number of reviews relating to ash slagging characteristics have already been reported. For instance, the advanced analytical method of slag constituents in coal using Computer-Controlled Scanning Electron Microscopy (CCSEM) is useful for the discussion on chemical aspect of the ash deposition [2,3]. Raask [4] elucidated deposit initiation, and Walsh et al. [5,6] studied deposit characteristics and growth. ⇑ Corresponding author. Tel.: +86 451 86418854; fax: +86 451 86412528. E-mail address: [email protected] (Z. Li). 0306-2619/$ - see front matter Ó 2010 Elsevier Ltd. All rights reserved. doi:10.1016/j.apenergy.2010.08.015

Beer et al. [7] tried to develop modeling ash deposition. Vuthaluru et al. [8] have studied formation of ash intermediates that consist of gases, liquids and solids. Benson et al. [9] summarized the behavior of ash formation and deposition in coal combustion. Naruse et al. [10] performed experiments on ash deposition based on a horizontal pulverized coal reactor with a pre-combustor. Even from these references, however, precise and quantitative understanding of slagging characteristics during coal combustion or gasification at high temperature has still been lacking because slagging strongly depends on the structural shape of the burner. Consequently, it is desirable to study the influence on slagging characteristics of key parameters pertinent to the structure of the burner. The use of computational fluid dynamics (CFD) models constitutes a powerful tool to study and characterize some complex processes that take place in the boiler. CFD provides a great deal of precise numerical calculations of velocity, temperature and concentration fields, irradiation profiles, heat transfer distribution and pollutant formation. Sarlej et al. [11] demonstrated an application of CFD in experimental burner design. Taglia et al. [12] presented and discussed optimization studies based on computer simulations of a 450 kW combustion chamber. Habib et al. aimed at conducting a numerical investigation of combustion parameter influences on NOx production in an industrial boiler. The CFD package FLUENT was used for the calculations in the study. Their CFD model predictions were in good agreement with experimental results [13]. CFD modeling has been applied to simulate a number of large-scale combustion facilities [14–17]. Commercial CFD codes such as FLUENT have also proven their worth in looking at control operations of and analyzing processes within coal-fired utility boilers [18–20].

Z. Li et al. / Applied Energy 88 (2011) 650–658

651

Nomenclature

l Pi

lref Tps NBO/T

a lH lL BH

BL

AH

particle viscosity (in Pa s) sticking probability of the particle group i with average viscosity l critical viscosity (in Pa s) temperature of particle (in K) ratio of nonbridging oxygens to tetrahedral oxygens (see Eq. (2)) ratio of mole fraction network modifiers to the sum of network modifiers and amphoteric (see Eq. (3)) high temperature viscosity (in Pa s) low temperature viscosity (in Pa s) high temperature viscosity parameter which adopts Senior and Srinivasachar’s coefficient for calculation (see Eqs. (4) and (5)) low temperature viscosity parameter which adopts Kalmanovitch and Frank’s coefficient for calculation (see Eqs. (6) and (7)) high temperature viscosity parameter which adopts the parameter BH for calculation (see Eq. (8))

For the present work, FLUENT was used to simulate the reacting flow field, based on these results, attention has been focused on slagging characteristics by adding various subroutines to handle the numerical modeling of slagging. Specifically, using computer simulations the slagging characteristics were investigated near the throat of the burner. The distribution of coal particles on slagging characteristic in the primary air ducts of different types of burners has been previously studied in Ref. [20]. This paper presents our studies on the influence of swirl number and burner structure on the sticking behavior near the burner throat. The motivation is the belief that particle sticking behavior can reflect the slagging condition within the burner. 2. Burner introduction A low-NOx axial-swirl coal burner [21] in a 600-MWe utility boiler designed by Mitsui Babcock Energy Limited was investigated (see Fig. 1 for a schematic diagram). The cross-sectional profiles show the partitioning of the air stream from wall to center within the burner associated with the outer secondary air, inner secondary air, and primary air. Swirling of the inner and outer secondary air begins after the air passes through the axial-register vanes. The swirling intensity of these streams can be regulated by changing the angle of the axial-register vanes. In contrast, the primary air/coal mixture begins swirling through the bent vanes in the primary air duct. After the swirling coal/air mixture enters the four axially-arranged channels, swirling is restrained and weakens. Coal inertia then gathers the pulverized coal near the collector and four fuel-rich and fuel-lean air/coal mixing zones form; thereby, circular fuel-biased combustion is achieved. Four pulverized coal collectors are installed in the primary air duct. Under actual operating conditions, the inner and outer secondary air vanes are set at 50° and 25°, respectively, with regard to burner axis. The design parameters of the burner are given in Table 1. An analysis of coal and ash is presented in Table 2. Fig. 2a and b present photos of the burner after developing slagging problems during boiler operations, with the first showing burn damage and distortion and the second highlighting the point that slagging near the water-cooled wall had accumulated mainly around the burner throat. At distances away from the burner throat the quantity of slag decreases significantly.

AL S0 Gh Gx R U W p

c n A

g N N0

low temperature viscosity parameter which adopts the parameter BL for calculation (see Eq. (9)) swirl number axial flux of the tangential momentum axial flux of the axial momentum outer radius of the annulus. axial components of the velocity tangential components of the velocity static pressure sticking number density number of sticking particles statistical slagging wall area (in m2) sticking-particle ratio total number of sticking particles number of tracked particles, which in our simulations is N0 = 960,000

3. Mathematical model and calculation method 3.1. Mathematical model of the burner flow field Using the commercially-available FLUENT 6.3.26 software, the flow field of a burner can be calculated with a variety of widelyused numerical models. Gas turbulence was specifically taken into account by the so-called realizable k–e model [22]. The Lagrangian stochastic tracking model was applied to analyze the gas/particle flow field [23], while calculations of the gas/particle two-phase coupling use the Particle-Source-In-Cell (PSIC) method [24]. Radiation was described using the P-1 model [25], and devolatilization was modeled by a two-competing-rate Kobayashi model [26]. The combustion of volatiles was modeled using Probability Density Function theory [26], and char combustion by the diffusion/kinetics model [27]. 3.2. Slagging model of the water-cooled tube wall To calculate slagging characteristics near the water-cooled tube wall, FLUENT software was used to compute the reacting flow field of the single burner, and then added a slagging subroutine to FLUENT based on results of the reacting flow field. The subroutine was written in the C language, with user-defined functions. During simulations, particles striking the water-cooled tube walls triggered calls to this subroutine. Fig. 3 shows a flow chart of the subroutine procedures involved in calculating the sticking behavior of these particles. The slagging computation is comprised of three main elements. The particle viscosity is calculated first, and the particle sticking probability is then computed using this viscosity. With the particle sticking probability, the number of sticking particles is calculated. The particle viscosity is the viscosity of ash particles melting from solid to liquid under high temperature and is calculated by the following procedure [28–32]: (i) Determine the mole fractions of all components based on the oxide composition of ash:

NBO=T ¼ ðCaO þ MgO þ FeO þ Na2 O þ K2 O  A12 O3 Þ=ððSiO2 þ TiO2 Þ=2 þ A12 O3 Þ:

ð1Þ

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Z. Li et al. / Applied Energy 88 (2011) 650–658

Fig. 1. Schematic of the low-NOx axial-swirl burner: (1) central air entrance, (2) inner secondary air duct, (3) outer secondary air duct, (4) primary air entrance, and (5) pulverized coal collector.

(iii) If BH < 0 or BH > 50.0, use Kalmanovitch and Frank’s coefficients to calculate BH, where

Table 1 Design parameters of the low-NOx axial-swirl burner in the utility boiler. 2

Exit area of the central air (m ) Exit area of the primary air (m2) Exit area of the inner secondary air (m2) Exit area of the outer secondary air (m2) Temperature of the central air(K) Temperature of the primary air (K) Temperature of the secondary air (K) Mass flow rate of the central air (kg s1) Mass flow rate of the primary air (kg s1) Mass flow rate of the inner secondary air (kg s1) Mass flow rate of the outer secondary air (kg s1) Fuel consumption (kg s1)

BH ¼ BK0 þ BK1  a þ BK2  a2 þ SiO2  ðBK3 þ BK4  a 0.0585 0.1628 0.2278 0.5908 599.6 351 599.6 0.61 4.98 2.32 13.09 2.97

þ BK5  a2 Þ þ ðSiO2 Þ2  ðBK6 þ BK7  a þ BK8  a2 Þ þ ðSiO2 Þ3  ðBK9 þ BK10  a þ BK11  a2 Þ:

ð4Þ

(iv) Calculate the low temperature BL using Senior and Srinivasachar’s coefficients where

BL ¼ BL0 þ BL1  a þ BL2  a2 þ SiO2  ðBL3 þ BL4  a þ BL5  a2 Þ þ ðSiO2 Þ2  ðBL6 þ BL7  a þ BL8  a2 Þ þ ðSiO2 Þ3  ðBL9 þ BL10  a þ BL11  a2 Þ:

ð5Þ

(v) If BL < 10.0 or BL > 60.0, use modified coefficients to calculate BL where

BL ¼ BW0 þ BW1  a þ BW2  a2 þ SiO2  ðBW3 þ BW4  a

Table 2 Coal characteristics and ash analysis.

þ BW5  a2 Þ þ ðSiO2 Þ2  ðBW6 þ BW7  a þ BW8  a2 Þ

Proximate analysis (as received, wt.%) Fixed Ash Moisture Volatiles carbon

Net heating value (kJ kg1)

50.53

28.15

24,450

Nitrogen 0.64

Oxygen 9.76

7.39

13. 93

Ultimate analysis (as received, wt.%) Carbon Hydrogen Sulfur 64.33 3.65 0.3

þ ðSiO2 Þ3  ðBW9 þ BW10  a þ BW11  a2 Þ:

ð6Þ

(vi) Calculate the high temperature AH where

AH ¼ 2:81629  0:46341  BH  0:35342  NBO=T:

ð7Þ

(vii) Calculate the low temperature AL where

AL ¼ 0:982  0:902473  BL if NBO=T P 1:3: Ash analysis Ash analysis (wt.%) SiO2 Al2O3 Fe2O3 36.29 21.60 11.57

AL ¼ 2:478718  0:902473  BL  2:662091  NBO=T Na2O 0.91

K2O 0.99

TiO2 0.97

Ash fusibility temperatures (K) Deformation Softening temperature temperature 1443 1533

CaO 16.64

MgO 1.22

SO3 9.56

MnO2 0.26

if 0:2 6 NBO=T < 1:3: AL ¼ 9:223  0:902473  BL  36:3835  NBO=T

Fusion temperature

if 0:0 6 NBO=T < 0:2: AL ¼ 9:223  0:902473  BL if NBO=T < 0:0:

1563

(viii) Calculate the high temperature viscosity lH (Pa s) at a given temperature T (K) where

a ¼ ðCaO þ MgO þ FeO þ Na2 O þ K2 O þ 2  TiO2 Þ=ðCaO þ MgO þ FeO þ Na2 O þ K2 O þ 2  TiO2 þ A12 O3 Þ:

ð2Þ

(ii) Calculate the high temperature BH using Senior and Srinivasachar’s coefficients where

BH ¼ BH0 þ BH1  a þ BH2  a2 þ SiO2  ðBH3 þ BH4  a

lH ¼ T  10AH þ 1000  BH=T:

ð3Þ

ð9Þ

(ix) Calculate the low temperature viscosity lL (Pa s) at a given temperature T (K) where

lL ¼ T  10AL þ 1000  BL=T:

þ BH5  a2Þ þ ðSiO2 Þ2  ðBH6 þ BH7  a þ BH8  a2 Þ þ ðSiO2 Þ3  ðBH9 þ BH10  a þ BH11  a2 Þ:

ð8Þ

ð10Þ

The larger viscosity value from the high and the low temperatures is selected as the particle viscosity.

Z. Li et al. / Applied Energy 88 (2011) 650–658

(a)

653

(b)

Fig. 2. Deformation of the burner ports and slagging near the burner throat: (a) deformation of the burner ports and (b) slagging near the burner throat.

Fig. 3. A flow chart outlining the procedures in calculating particle sticking behavior.

l ¼ maximum of ðlH; lLÞ:

ð11Þ

Particle temperature T is determined by the temperature field, which is calculated by the combustion model, and read by the macro function P(T). The Senior and Srinivasachar coefficients BH0–BH11 and BK1– BK1 and the Kalmanovitch and Frank coefficients BL0–BL11 and BW0–BW11 can be found in Ref. [32]. Particle sticking probability is calculated as follows. If the particle viscosity is smaller than the critical viscosity, the particle can be considered to have stuck completely to the water-cooled tube wall, upon which the sticking probability is taken as 1; if the particle viscosity is otherwise larger than the critical viscosity, the sticking probability is then defined as the ratio of the critical and particle viscosities. The relationships are summarized in Eq. (12) [33]:

Pi ðT ps Þ ¼ lref =l Pi ðT ps Þ ¼ 1

l > lref :

l 6 lref :

ð12Þ

where Pi is the sticking probability of the particle group i with average viscosity l, Tps is the temperature of particle group i, and lref is the critical viscosity. Generally, when slag changes from the liquid to the plastic state, there is a transitional point in the viscosity curve. At this temperature, a large number of crystals vanish, and the viscosity at the transitional point is usually taken as the critical viscosity. The notable advantage of this model is that it is simple and appropriate for engineering applications. All current studies use this model to simulate the slagging process and to estimate the sticking-particle number. Generally, the critical viscosity is treated as constant during slagging, and the value commonly selected is 105 Pa s [32,34]. In the present study, a temperature subarea method [28] has been used to compute particle viscosity. The viscosity formulas for high and low temperatures are different, and the maximum value from these two viscosities is selected. This model is sufficiently precise if the particle critical viscosity is in the range 104–109 Pa s, which includes the commonly selected critical viscosity. Consequently, this model can give credible results and has wide application.

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n a review of slagging models, it was concluded that many factors influence this process – for example, particle temperature, ash fusibility temperatures and ash components, but without the factors of particle incident angle and velocity. With regard to the reacting flow field, a stochastic trajectory model was used to compute particle trajectories and to collect statistics on sticking-particle number. In total, particle trajectories numbered 960,000. 3.3. Computational domain specifications and calculated parameters The reacting flow of a full-size burner was simulated. The size of the reacting flow region was 7816 mm long, 10025 mm high and 7935 mm wide (see Fig. 4). The distance from the coordinate origin to the water-cooled tube wall was 550 mm. A schematic diagram of the computational domain and the origin of coordinates are shown in Fig. 4. A 3D mesh with 664,274 cells was allocated over the computational domain. A refined grid was constructed in the regions of maximum interest where a sudden change in fluid flow is expected, for example close to areas representing the burner exit and around the water-cooled tube walls of the slagging calculation. The finest grid with the greatest density of cells was constructed near the burner region. The operational parameters are listed in Table 1, while the coal proximate analysis and the ash analysis are given in Table 2. From a comparison of coal proximate analysis and ash analysis using the viscosity model improved by Wang [32] and coal proximate analysis and ash analysis using models adopted in our calculation, small differences in the ranges of moisture were found. The ranges of other terms are all within the ranges of the coal proximate analysis given in [32] and so it is reasonable to adopt this viscosity model. The average particle diameter of pulverized coal particles is 50 lm with particle sizes of 40–60 lm accounting for 70%, 0–40 lm accounting for 17%, and 60– 120 lm accounting for 13%. The proportion of pulverized coal particle with particle diameter close to the mean diameter was high. When pulverized coal is heated, coal particles soften and revert to a plastic state. Moreover, particle sizes become much smaller and particle shape becomes rounder as volatile is released. The following assumption that the pulverized coal particles are spherical [35–39] with a diameter of 50 lm was made.

Fig. 4. Schematic diagram of the computational domain (dimensions in mm).

3.4. Verification of the model A cold single-phase experimental test was initiated to investigate the flow field. This test employed a 1/4-sized model version of the full-scale burner with only air, and no particles, circulating in the low-NOx axial-swirl coal burner. An IFA300 constant-temperature anemometer system was used to measure the air velocity at the measurement points. Using a probe with hot-film sensors, the three-dimensional flow field at the exit of the burner was measured [40]. All measurement procedures are the same as reported in Ref. [41]. To avoid particles depositing on the film and causing damage during measurements, the room was kept clean and the air kept dry. Before each experiment, all measuring devices were calibrated. By repeating the experiment, measurement accuracy was assured [42]. The error for all velocity measurements was less than 5%. A three-dimensional computer model based on the burner model was configured for use in the single-phase numerical calculation. This configuration closely mimics the experimental rig. The grid division method and grid quantity were set-up to be the same as in the full-scale burner numerical calculation. The realizable k–e model was adopted to simulate the gas turbulent flow. Figs. 5–7 present comparisons of experimental and numerical results of the design burner. In these figures, x represents the axial distance to the exit of the burner along the jet flow direction, d the diameter of the outermost cone of the burner, which is 376 mm, and r the distance to the center line of the burner along the radial direction. The coordinate axes and origin are displayed in Fig. 1. Fig. 5 shows the gas axial velocities of both the test and the numerical simulation at x/d = 0, 0.25, 0.5, 1, 1.5 and 2.5. At x/ d = 0, axial velocities are negative in the central zone indicating that the recirculation zone of burner is generated earlier and has extended into the burner throat. The unburned pulverized coal and ash particles entrained by high temperature gas collide continually with the burner throat. This can easily cause slagging and deformation of the burner, a phenomenon that can occur under normal operations. At x/d = 1.0, negative velocity values disappear. At x/d = 0, axial velocities have two peaks; the one nearest the burner center is in the inner secondary air flow area while the other is in the outer secondary air flow area. As x/d increases, primary air disperses into the secondary air which is pushed towards the side wall, and the peak value of the axial velocities decreases gradually. Fig. 6 shows the distribution of radial gas velocities. From x/d = 0 to 0.5, there is an obvious peak zone at r P 150 mm and lies within the zone of secondary air flow. Between x/d = 0 and 0.25, there is a larger radial velocity zone. With increasing x/d, the maximum radial velocity decreases gradually. Radial velocities are negative in the outer zone because of the outer recirculation flow. Fig. 7 shows the distribution of tangential gas velocities. Between x/d = 0 and 0.5, with the influence of tangential secondary air, there is an obvious peak zone at r P 150 mm. Peak values decrease with the diffusion of gas flow. Experimental measurements coincide well with the simulation from x/d = 0 to 0.5. From x/d = 1.0 to 2.5, experimental tangential velocities decrease quickly falling to zero at x/d = 2.5. However, tangential velocities are still in good agreement with simulations. From the above comparison, axial and radial velocities from simulations agree well with the experiment, although tangential velocities from simulations decrease more slowly than those from experimental measurements especially in the range x/d = 1.0–2.5. The main reason for this is that the realizable k–e model is adapted to simulate isotropic turbulence, but the real flow is anisotropic. As a consequence, results from simulations can basically show only the distribution of gas velocities.

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Axial velocity (m/s) -5 0 5 10 15 -5 0 5 10 15 0

600

5 10 15 0

5

10

0

5

10

0

5

10

500

Radius (mm)

400

300

200

100

0 x=0 x/d=0

x=373 mm x=559.5 mm x/d=1.5 x/d=1.0 − calculated value

x=93.25mm x=186.5 mm x/d=0.25 x/d=0.5 experiment value

x=932.5 mm x/d=2.5

Fig. 5. Distribution of mean gas axial velocities in the test and the numerical simulation.

Radial velocity (m/s) -5 600

0

5

10

0

5

10

0

5

0

5

0

2

4

0

2

4

Radius (mm)

500

400

300

200

100

0 x=0 x/d=0

x=93.25mm x=186.5 mm x/d=0.25 x/d=0.5 experiment value

x=373 mm x=559.5 mm x/d=1.0 x/d=1.5 − calculated value

x=932.5 mm x/d=2.5

Fig. 6. Distribution of mean gas radial velocities in the test and the numerical simulation.

4. Computational results and analysis Twenty walls equally distributed above and below the burner center were introduced. The height of each wall was 70 mm (equal to 0.047D where D = 1504 mm is the diameter of the burner’s outermost port). Fig. 4 shows the arrangement of these walls centered equally about the burner’s outermost port. The sticking number density is defined as the ratio of the number of sticking particles to the wall area as summarized in Eq. (13):

c ¼ n=A;

ð13Þ

where c is the sticking number density, n is the number of sticking particles, and A is the wall area, m2. The coal feeding flux was unchanged for the different settings under study, a listing of which can be found in Table 3. 4.1. The slagging condition of the designed burner Fig. 8 shows the sticking number density distribution calculated along the water-cooled wall of the designed burner. The slagging condition at Y  D/2 > 0 is more pronounced than that at Y  D/2 < 0. Over the water-cooled wall at Y  D/2 > 0 and Y  D/2 < 0, the sticking number density first increases and then decreases with

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Tangential velocity (m/s) -10 -5 0 5 10 -5 0 5 10 15 0 5 10 15 600

0

5

10

0

10 0

5

5

10

Radius (mm)

500 400 300 200 100 0 x=0 x/d=0

x=93.25mm x=186.5 mm x=373 mm x=559.5 mm x=932.5 mm x/d=0.25 x/d=0.5 x/d=1.0 x/d=1.5 x/d=2.5 experiment value − calculated value

Sticking number density [1/(m2)]

Fig. 7. Distribution of mean gas tangential velocities in the test and the numerical simulation.

Table 3 Slagging computation settings. Changes in the vane angle of the outer secondary air Vane angles of the outer secondary 10° 15° air Swirl number 0.09 0.117

25°

35°

45°

0.126

0.294

0.414

Changes in the proportion of inner and outer secondary air Mass flow rate of the inner secondary 1.86 2.32 air (kg s1) 13.55 13.09 Mass flow rate of the outer secondary air (kg s1) Swirl number 0.12 0.126 Changes in the dimensions of the burner ports Distance (mm) Structure 1 (design structure) L1 480 L2 371 L3 268

2.78

3.25

12.63

12.16

0.128

Structure 2 480 371 0

a dependency on the absolute value of Y  D/2, with a maximum value at |Y  D/2| = 105 mm. The calculated total particle sticking number was 318522 which accounts for 33% of the total number of particles tracked and represents in this case a large and serious slagging problem for the designed burner. Thus, the results from slagging simulations were basically consistent with results obtained under normal operating conditions (see Fig. 2b). 4.2. Influence of the secondary air change on slagging Swirl number is an important parameter characterizing the combustion process in furnace. As was originally proposed by Chigier and Beer [43], the degree of swirl for a swirling flow is usually characterized by the swirl number:

S0 ¼

Gh ; RGx

5x104 4x104 3x104 2x104 1x104 0 0

0.132

Structure 3 212 103 0

where Gh stands for the axial flux of the tangential momentum, Gx is the axial flux of the axial momentum, and R is the outer radius of

140

280

420

560

700

The absolute value of Y-D/2 (mm) Fig. 8. Sticking-particle number density distribution along the water-cooled wall in the designed burner from numerical calculation.

the annulus. These two parameters, Gh and Gx can be expressed as [43,44]:

Gh ¼

Z

R

ðWrÞqU2pr dr;

ð15Þ

0

Gx ¼

Z

R

2prqU 2 dr þ

Z

0

R

2prp dr;

ð16Þ

0

where U, W, and p are the axial and tangential components of the velocity and static pressure, respectively. It is very difficult to directly estimate the pressure integral term because static pressures are strongly dependent on the geometry of the swirl. According to Beer and Chigier [45], Martin [46], and Weber and Dugue [47], the axial momentum flux, Gx, can be obtained with good approximation by eliminating the pressure term in Eq. (16). Thus, the modified swirl number S defined in Eq. (14) can generally be expressed as:

S0 ¼ ð14Þ

Y-D/2>0 Y-D/2<0

6x104

Gh

2pR

RR 0

qU 2 r dr

RR ¼

qUWr2 dr : R 0 qU 2 r dr 0

RR

ð17Þ

With constant total secondary air mass flux, the swirl number of the burner is varied by changing the angle of the axial-register vanes in the outer secondary air duct and the proportion of the

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30 25 20 15 10

50

Sticking particle ratio (%)

(b)

35

Sticking particle ratio (%)

(a)

40 30 20 10

5 0.10 0.15 0.20 0.25 0.30 0.35 0.40

0.120

0.125

0.130

0.135

Swirl number

Swirl number

Fig. 9. The influence of changing the calculated swirl number on particle sticking rate: (a) changes in blade angle of secondary air and (b) changes in proportion of the inner secondary air and the outer secondary air.

ð18Þ

where g is the sticking-particle ratio, N is the total number of sticking particles, N0 is the number of tracked particles, which in our simulations is N0 = 960,000. Fig. 9a shows the influence of the angle of the axial-register vanes in the outer secondary air duct obtained from numerical calculations. With the angle of axial-register vanes set at the design angle of 25°, the swirl number is 0.126 with a sticking-particle ratio of 33.2%. At swirl numbers of 0.09, 0.117 and 0.414, the sticking-particle ratio is lower. If the swirl number is smaller than 0.126, then as the swirl number decreases, decreases in the rotation of the flow, the number of the pulverized coal particles colliding the water walls, and the sticking-particle ratio were found. If the swirl number is larger than 0.126, then, as swirl number increases and the rotation of the flow gets stronger, mixing between the high temperature gas and the pulverized coal particles becomes greater, thereby promoting the ignition and burnout of the pulverized coal particles. In turn, the particle concentration near the burner decreases, the particle number colliding with the water-cooled tube walls decreases and the sticking-particle ratio also decreases. The relationship between swirl number and sticking-particle ratio is non-linear. In particular, at swirl number of 0.117, the sticking-particle ratio is at a minimum of 8.73% and therefore giving an optimized angle of 15°. Fig. 9b shows the effect of the proportion of the inner and outer secondary air mass flux as obtained from calculations. This proportion has less effect on swirl number, but it has a greater impact on sticking-particle ratio. As the swirl number increases, the stickingparticle ratio decreases. Meanwhile, as secondary air increases, oxygen needed for coal burning increases in the burner central zone. As a consequence, coal particles burn faster and the number of the particles impinging on the wall then decreases. By increasing the mass flux of the inner secondary air from the design value of 2.32–2.78 kg s1, the swirl number increases from 0.126 to 0.128, and the sticking-particle ratio falls from 33.2% to 11.8%. This decrease is more pronounced because a further increase in the secondary air to 3.25 kg s1 elicits an increase in the swirl number to 0.132, while decreasing the sticking rate to the marginally lower value of 9.6%.

35

Sticking particle ratio (%)

g ¼ N=N0 :

been made (see Table 3). Fig. 10 shows slagging conditions for these different-sized burner ports obtained from these simulations. The sticking-particle ratio is largest for the prototype structure. The burner port with the least sticking-particle ratio, i.e., 18.3%, is structure 2 with L3 = 0 mm while structure 1 has the highest sticking-particle ratio. Fig. 11 presents axial velocity distributions at the burner outlet indicating negative values in the burner central zone for structure 1. The high temperature gas of the central recirculation zone has extended into the burner throat resulting in the burner ports distorting easily, and the unburnt coal particles are easily thrown toward the water-cooled walls. The total number of the particles colliding with the walls as well as of those sticking to the walls then increases. While the axial velocities for structures

30 25 20 15 10 1

2

3

The burner structure Fig. 10. Influence of burner structure on slagging from numerical calculation.

40

Axial velocity (m/s)

inner and outer secondary air mass flux (see Table 3). The stickingparticle ratio is defined as the ratio of the total number of sticking particles to the number of tracked particles, that is:

structure 1 structure 2 structure 3

30 20 10 0 -10

4.3. Influence of burner port structure on slagging To resolve the heavy slagging and deformation problems within the burner port, simulations with differing port dimensions have

-1.5

-1.0

-0.5

0.0

0.5

1.0

1.5

r/R Fig. 11. Axial velocity distributions at the burner outlet.

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Z. Li et al. / Applied Energy 88 (2011) 650–658

2 and 3 are positive, there is no central recirculation zone in the burner throat. The sticking-particle ratio of structure 2 is less than for structure 3. It is clearly advantageous to protect the burner ports and in this regard structure 2 provides an optimal solution in reducing slagging. 5. Conclusions Numerical simulations and experiments of single-phase flows and numerical simulations of slagging occurring in a single swirl coal combustion burner of a 600-MWe bituminous coal-fired boiler were performed. The following conclusions have been obtained: A comparison of numerical simulations and experiments of single-phase flows indicate that the mathematical model employed in the simulation have given reasonable quantitative accounts slagging characteristics. The relationship between outer secondary air vane angle and sticking-particle ratio is non-linear. With an outer secondary air vane angle of 25°, the sticking-particle ratio is as much as 33.2%. As the outer secondary air vane angle was varied either side of 25°, the sticking-particle ratio decreased. With an outer secondary air vane angle of 15°, the sticking-particle ratio was at its lowest value of 8.73%. As the mass flux of the inner secondary air was increased, the swirl number increased while the sticking-particle ratio decreased. Structure 2 yielded the lowest sticking-particle ratio, i.e. 18.3%, and hence found to hold the optimal values of the three structures.

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