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Erosion of heat exchangers due to sootblowing
Engineering Failure Analysis 33 (2013) 473–489
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Erosion of heat exchangers due to sootblowing Wacław Wojnar ⇑ Silesian University of Technology, Faculty of Energy and Environmental Engineering, Institute of Power Engineering and Turbomachinery, Division of Boilers and Steam Generators, Konarskiego 20, 44-100 Gliwice, Poland
a r t i c l e
i n f o
Article history: Received 20 May 2013 Received in revised form 21 June 2013 Accepted 22 June 2013 Available online 4 July 2013 Keywords: Erosion Superheaters Sootblower CFD
a b s t r a c t The erosion rate of convection surfaces is crucial for boiler reliability. This paper presents the impact of the action of a sootblowing stream (steam and air) jet on the sustainability of tubes. The model investigations described here compared the erosion rates of a tube bank while turning on or off an additional jet stream of compressed air that simulates the operation of a sootblower. The numerical model code Fluent was used. Based on mass loss measurements and numerical calculations, it can be confirmed that the gas velocity has the greatest impact on the particulate erosion. This velocity depends on the stagnation parameters at the inlet of the nozzle. This paper presents calculations for the erosion loss of a heat exchanger built in the actual object. Ó 2013 Elsevier Ltd. All rights reserved.
1. Introduction In boilers that fire solid fuels, various erosion processes can occur, leading to serious operational and economic problems. The wear of the heat exchanger tubes depends, among other factors, on: The velocity of solid particles, which is related to the erosion rate as a power function. The experimentally determined powers have been compared [1], and they amount to 2.4–3.5 for the wear of metals, approximately 3 for ceramics and approximately 5 for polymers. The concentration of the solid phase in the gas. At low concentrations, the erosion rate increases with the concentration. When the concentrations are high, the influence of this parameter may be insignificant or even decrease the wear observed due to collisions between the incident and reflected particles. The particle diameter: the erosion rate increases until the particle diameter reaches the range of 50–100 lm. Above this range, the erosion rate is not influenced by the particle diameter. The shape of particles: for metallic surfaces, non-spherical, sharply edged particles cause stronger wear. The hardness of the erodent. The impact angle ue between the two-phase stream and the surface: the maximum wear of ductile materials occurs for ue max = 20°–40° [2,3], while that of brittle materials occurs for perpendicular impact. The erosion rates of the convection surfaces are crucial for the boiler reliability. Many papers, for example [1–8], have described tube failures resulting from wear due to the fly ash in the flue gas, but the erosion caused by particulate matter entrained in the sootblowing stream (steam or air) has not been investigated at the same level. Because the existing methods
⇑ Tel.: +48 322372374; fax: +48 32 2372193. E-mail address: [email protected] 1350-6307/$ - see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.engfailanal.2013.06.026
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Nomenclature Ar Cx c cbl cmax da da min da max dbl dl D ee Fl FL Fr Fu g i i0 iL ma m_ a _ bl m mst n p P p0 p1 pam pc pL Re Rea s s1 s2 t0 y z1 z2 ve = Dh/s ve bl = Dhbl/sbl-a V V_ V_ n w w2 wa waf wg win wL wm wmax w wbl
Archimedes number współczynnik oporów ruchu ziarna pyłu w gazie, mean concentration of fly ash in the flue gas (sootblower off), g/m3 mean concentration of fly ash in the flue gas (sootblower on), g/m3 maximum concentration of fly ash in the flue gas, g/m3 particle diameter, lm minimum particle diameter, lm maximum particle diameter, lm blower nozzle diameter, m inside diameter of the blower lance, m outer tube diameter, m erosion coefficient defining the erosive properties of the fly ash, lm/h blower lance cross-section, m2 total steam outflow area, m2 resistance force of the spherical particle with diameter da, N uplift force of the gaseous medium with the density qg, N acceleration of gravity, m/s2 enthalpy, kJ/kg stagnation enthalpy, kJ/kg steam enthalpy at the blower nozzle outlet, kJ/kg solid phase mass (erodent mass), kg, erodent stream, kg/s steam mass flow from the blower, kg/s steam mass per 1 blowing cycle, kg polydispersion index steam pressure at the blower inlet, bar gravity force of solid phase mass ma, N stagnation pressure, Pa pressure on the inlet of the nozzle, bar, ambient pressure, bar, pressure on the outlet of the compressor, bar steam pressure at the blower nozzle outlet, bar Reynolds number, Reynolds number for particle fall entropy, kJ/kgK transverse pitch, mm longitudinal pitch, mm stagnation temperature, °C height coordinate, mm number of tubes in the first row number of rows, erosion rate, lm/h blower operation-related erosion rate, lm/h grid volume, m3 supply fan output, kg/s, m3/s air stream in the nozzle, kg/s steam flow velocity in the blower lance, m/s blowing agent velocity outside the nozzle, m/s particle velocity, m/s fall velocity, m/s flue gas velocity in the free section at the first tube, m/s inlet velocity of the test duct (level y = 700 mm, Fig. 4b), m/s steam velocity at the blower nozzle outlet, m/s velocity in the cross-section of the nozzle (Fig. 2), m/s maximum velocity of the dust-laden flue gases in the duct section, m/s mean flue gas velocity in the empty duct at the bank inlet (sootblower off), m/s mean flue gas velocity in the empty duct at the bank inlet (sootblower on), m/s
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475
Greek symbols:
amax b2 bc bL bm bw dbl Dh Dhbl Dmtbl Dmt Dp
gp lg q qa qg qL qt
r1 ¼ s1 =D r2 ¼ s2 =D s sn bl a sc bl sa sbla u ue ue max
angle spanning over eroded tube perimeter parameter defining the impact of longitudinal pitch of the bank on the erosive wear of the tubes parameter defining the fly ash concentration distribution in the flue gases critical pressure ratio parameter defining the impact of the operating temperature and the steel grade on the erosive wear parameter defining the flue gas velocity distribution in the boiler duct factor introducing the sootblowing effect thickness loss of the tube, lm, mm thickness loss of the tube (sootblower on), lm, mm test tube mass loss at air jet switched on (g) test tube mass loss at air jet switched off (g) differential pressure – orifice plate, Pa mean probability of ash particle impingement on the tube gas dynamic viscosity, kg/m s density, kg/m3 erodent density, kg/m3 gas density, kg/m3 steam density at the blower nozzle outlet, kg/m3 tube density, kg/m3 relative transverse pitch of the tube, m relative longitudinal pitch of the tube, m time interval, s annual time with the blower shut down, h blowing cycle duration time, s annual boiler operating time, h annual duration time of the erosion process caused by ash blowing, h air humidity, % angle between two-phase stream and the surface tube, ° maximum angle between two-phase stream and the surface tube, °
of calculation are not appropriate for this case, this work is an attempt to determine this phenomenon quantitatively to reduce the erosive damage (Fig. 1) by controlling the process parameters.
2. Calculation of the erosion rate during sootblowing The fly ash erosion rate (lm/h) of the plain tube banks in a flue gas stream may be described, according to [2,9], as
3 b3w ee cbc bm b2 gp Dh=s ¼ v e ¼ 2:33 107 w
ðlm=hÞ
cmax where bw ¼ wmax ; bc ¼ w c The method to calculate the parameters in Eq. (1) is presented in [10].
Fig. 1. Examples of erosion damage to convection surfaces.
ð1Þ
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
(a)
(b)
Tube bank
w
Flue gas with ash particles
wbl
wmax Sootblower nozzle
w max Sootblower nozzle
wm
¼ 5—10 m=s;wmax 1:25 wÞ and (b) the blowing jest turned on (w bl ffi w, Fig. 2. Scheme for the erosion rate calculation: (a) the blowing jest turned off ðw wmax = f(static parameters of the blowing medium p0, t0; distance nozzle-tubes; wm), wm – velocity in the cross-section of the nozzle, which is ca. 300 m/s for air and ca. 600 m/s for steam).
The problem investigated in this work differs from typical particle erosion problems on convective sections of steam boilers. An additional element is the jet stream of a high velocity soot-blowing medium, which entrains and accelerates the ash particles, directing them towards the tubes. The model investigations described below compare the erosion rates of a tube bank while turning the additional jet stream of the compressed air on or off, thereby simulating the sootblower operation. Assuming equal periods of rig operation in both states, the factor describing the sootblower effect is introduced as
dbl ¼
Dmtbl Dmt s¼idem
ð2Þ
Because the basic formula (1) describes the erosion loss as a reduction of the tube wall thickness in the area of the most intense erosion, the same formula can be used to determine sootblowing-induced erosion (3) for the case when the sootblowing nozzle is turned on. Both cases are presented in Fig. 2.
Dhbl
sbl a
3 b3w ee cbc bm b2 gp Þ ¼ v ebl ¼ 2:33 107 ðw bl
ðlm=hÞ
ð3Þ
As shown in Fig. 6, during sootblowing, the area of the most intense wear on the tubes occurs at an angle of approximately 40° symmetric about the inflow axis. Many investigations have shown that the maximum wear at this location also occurs during typical fly ash erosion without blowing. Assuming that the thickness loss Dh occurs along the radius R (from R1 to R2)
Fig. 3. Indication of tube wall thickness losses.
W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
(a)
477
(b)
Fig. 4. (a) Layout of the test stand: 1 – tube bank, 2 – test duct, 3 – fan, 4 – erodent container, 5 – feeder, 6 – cyclone precipitator, 7 – compressor, 8 – orifice plate, 9 – rubber tube, 10 – manometer, 11 – nozzle, 12 – orifice plate and (b) test duct with traverses.
and that the area of the most intense wear is located at the section of the perimeter described by the angle amax (Fig. 3), the following equation can be used:
dbl ¼
Dmt bl DV tbl qt amax ðR21 R22 Þtbl Dhbl ð2R1 Dhbl Þ ¼ ¼ ¼ Dmt s¼idem DV t q t Dhð2R1 DhÞ amax ðR21 R22 Þt
ð4Þ
while R1 Dh, this equation becomes
dbl ¼
Dmt bl Dhbl Dmt s¼idem Dh
Eventually, it can be assumed that the factor dbl
dbl ¼
Dhbl Dmt bl ¼ Dh s¼idem Dmt s¼idem
ð5Þ
proves that the mass loss measurements is equal to erosion rate and reflects the tube wall thickness loss for a given time period.
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2.1. Test rig and the method of investigation At the Institute of Power Engineering and Turbomachinery of the Silesian University of Technology, there is a test stand for investigating the phenomena of erosion, which is shown in Fig. 4 [1,2,11]. The test stand consists of a fly erodent container (4), a feeder (5), a duct connected to a supply fan (3), a test duct with the tube bank under investigation (1), a cyclone precipitator that (6) recirculates the erodent to the container and an output duct with a stack. The compressed air from two compressors is directed to the test duct (1) below the tube bank using a nozzle (11) with a diameter of dbl = 4 mm located exactly on the axis of the test duct (2). The air stream was measured with an ISA orifice plate (8). The differential pressure on the orifice plate was measured with a U-tube gauge. The air temperatures were measured using mercury thermometers. A pressure gauge (10) measuring the static pressure p1 at its inlet was fixed in front of the nozzle. The measured values of p1 were later compared with those calculated by Fluent [12]. The air stream at the outlet of the supply fan was also measured with an orifice plate (12), as shown in Fig. 4a. The supply air stream was tuned using the control flap located on the fan inlet to reach values in the range of 6–10 m/s at the cross-section of the first row of the tube bank, typical for coal fired boilers. The model steel tubes were covered with blue (outer layer), yellow and red (inner layer) paint according to the idea in [13]. In Fig. 4b, the following traverses (levels according to height – y coordinate) are shown:
y = 320 mm y = 500 mm y = 650 mm y = 715 mm
– – – –
compressed air nozzle outlet, cross-section where the velocity distribution was measured and modelled by Fluent, on this level, using Fluent, the velocities and particle concentrations were estimated to analyse eq. (1), cross-section of the first row of the tube bank.
The test tube bank arrangement was as follows:
external diameter of tubes D = 21.3 mm transverse pitch s1 = 67 mm longitudinal pitch s2 = 90 mm number of tubes in the first row z1 = 5 number of rows z2 = 8.
The air-erodent mixture circulates within this system. The fly ash from a PC boiler, which is described in Table 1, was used as the solid phase. Nearly 80% of the particles remained between 40 and 120 lm, with a polydispersion index of n = 0.912. It was decided that the similarity of the velocity, rather than the similarity of the Re number ðRe ¼ wg Dqg =lg Þ, is preserved in the test rig. In an actual boiler, Re = 2000 – 3000, and this value depends on the flue gas temperature and the velocity, which is in the range of wg = (6–9) m/s in PC boilers. The assumption of preserving the same Re value in the test stand kept at 0 °C, yields velocities equal to approximately 1 m/s, which makes pneumatic transport of the fly ash impossible. Consequently, the test rig is not capable of modelling all similarity criteria in parallel. Nevertheless, it is the velocity that is the most important parameter that influences the particulate erosion. Therefore, in the model duct, the velocities were kept exactly the same, as in an actual boiler. 2.1.1. Discussion of test results Three cases of modelling were analysed [14]: r s t
– – – – – –
air blowing nozzle ‘on’ _ a ¼ 0:06076 kg/s (high concentration) erodent stream m air blowing nozzle ‘on’ _ a ¼ 0:03383 kg=s (low concentration) erodent stream m air blowing nozzle ‘off’ _ a ¼ 0:03383 kg=s: erodent stream m
The operating data for the test stand required to estimate the representative compressed air stream are given in Table 2. Table 1 Characteristics of the erodent (fly ash). Density
Chemical composition (%)
qa
CaO 2.88
2700 kg/m3
MgO 2.41
Fe2O3 12.99
Na2O 0.51
Fineness Rx (%) K2O 1.74
SiO2 60.00
Al2O3 18.89
TiO2 0.57
R0.025 58.8
R0.032 52.2
R0.040 48.5
R0.063 36
R0.088 22
R0.12 11.3
R0.20 2.6
R0.30 0.6
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489 Table 2 Measurement results of the compressed air stream at the nozzle (dbl = 4 mm).
Temperature Pressure on the outlet of the compressor Pressure on the inlet of the nozzle Differential pressure – orifice plate Air stream [14]
Symbol
Unit
Value
t pc p1 Dp V_n
°C bar bar Pa kg/s
8.0 3.7 3.0 2100 0.0092
Table 3 Test stand investigation results. Ambient parameters
r s t
Supply fan output
t
pb
u
°C
mm Hg
%
V_ kg/s
V_ m3/s
10 10 23
752 752 742
55 55 68
0.269 0.269 0.258
0.216 0.216 0.216
Air stream in the nozzle V_n
Stream of the erodent m_ a
Mean concentration c
kg/s
Max. velocity in the axis of the nozzle (y = 500 mm) w500 max m/s
kg/s
g/m3
0.0092 0.0092 0
38 38 5,5
0.06076 0.03383 0.03383
276 154 157
In Table 3, the results concerning the three cases given above are presented, whereas Fig. 4 depicts the velocity distribution in the test duct at the level y = 500 mm, where the velocity probe (WTI) was introduced. Table 3 also shows the ash concentration for a cubic meter of the transported air. According to the investigation results presented in Fig. 5, with the compressed air nozzle switched off, the mean velocity = 5.2 m/s (y = 500 mm), which corresponds to a velocity in the range of wg = 7 m/s in the free section at the in the duct was w first tube row level (y = 715 mm). This velocity range is similar to that in actual boilers. The inlet velocity to the test duct (level y = 700 mm, Fig. 4b) was win = 10.8 m/s. On purpose the first measurement was conducted with the air blowing nozzle switched on and at a high erodent concentration and served as a basis for further comparisons. The purpose was to determine the shortest period after which the lowest (red) paint layer was exposed (the metallic tube surface was still hidden). The time noted was 30 min, and therefore, this period of time was chosen as a reference for further investigation. Exemplary mass losses of tubes after 30 min of exposure to erosion are presented in Tables 4 and 5. With the exception of the mass loss of the central tube, which is shown in Table 4, the mass losses of the tubes were very small. The majority of the results were within the range of the measurement errors, and thus, some losses in Table 4 are negative (mass increase). Therefore, only the central tube was chosen for further investigation. _ a ¼ 0:03383 kg=s). Fig. 6 presents the test tube bank at certain times (air blowing nozzle on and stream of the erodent m During the measurement with the air blowing nozzle switched off, which simulates the boiler duct when the sootblowers are not in operation, a noticeable mass loss of the paint covering the tubes (approximately 0.1 g) was recorded only after 23 h of the test run. The loss after 30 min (the reference period of time) was not measureable with satisfactory precision, and thus, it was necessary to assume a linear, time-dependent mass loss function. With this assumption, the mass loss for the central tube in the first row was Dmt = 0.002124 g/30 min). To determine the factor describing the influence of sootblowing on the erosion rate dbl, which is given in Eq. (3), the distribution of the velocity and the concentration of the erodent is necessary. Both values were calculated using Fluent [12].
8
Velocity [m/s]
7 6 5 4 3 2
1
2
3
4
5
Measuring point Fig. 5. Velocity distribution at the cross-section of the test duct (y = 500 mm) air blowing nozzle out – fan output V_ ¼ 0:216 m3 =s.
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Table 4 _ a ¼ 0:03383 kg=s (low concentration). Mass losses of test tubes for case 2 – air blowing nozzle on and stream of the erodent m Mass loss Dmt (g) after s = 30 min Row/column
A
B
C
D
E
8 7 6 5 4 3 2 1
0.0077 0.0064 0.0055 0.0112 0.0066 0.0029 0.0068 0.0081
0.0042 0.0064 0.0167 0.0098 0.035 0.0196 0.0137 0.0135
0.0018 0.0029 0.0133 0.015 0.0183 0.021 0.0254 0.2292
0.0075 0.0041 0.001 0.0002 0.0052 0.0049 0.0064 0.0042
0.0111 0.0043 0.0066 0.0077 0.0008 0.0072 0.0051 0.009
Table 5 _ a ¼ 0:06076kg=s (high concentration). Mass losses of test tubes for case 1 – air blowing nozzle on and stream of the erodent m Mass loss Dmt (g) after s = 30 min Row/column
A
B
C
D
E
8 7 6 5 4 3 2 1
0.004 0.0084 0.0114 0.0185 0.0144 0.0115 0.0141 0.005
0.0086 0.0165 0.017 0.0207 0.0323 0.0402 0.0441 0.0178
0.0144 0.0231 0.022 0.0256 0.0221 0.036 0.0291 0.3815
0.0142 0.0164 0.0236 0.0186 0.0144 0.0372 0.037 0.0184
0.0041 0.0084 0.0252 0.0081 0.0062 0.0044 0.0116 0.004
3. Numerical modelling of erosion When modelling erosion in Fluent [12], the description of the concentration is based on the residence time of the particulate matter in the unit of the grid volume V:
c ¼
_ as m V
ðkg=m3 Þ
ð6Þ
_a m Transforming Eq. (6) to c ¼ Fw , where F represents the area of the cross-section of the grid element, shows the connection a between the velocity of the particles and the concentration. The velocity of the particles depends on the gravitational force and may be influenced by the flow direction (upwards or downwards). The flow direction consequently increases or decreases the velocity of the particles, which determines the erosion rate. In the test stand, the air flows upwards, and, according to Stokes law, the particles uniformly accelerate until they reach constant terminal velocities at which the resistance force equals the gravity force of the particle minus the uplift force:
P ¼ Fr þ Fu
ð7Þ
where P ¼ ma g – gravity force of solid phase mass ma; F r ¼ 3plg da wa – resistance force of the spherical particle with diama eter da; F u ¼ m qa qg g – uplift force of the gaseous medium with the density qg. Because the particles are of diameters >10 lm, the Brownian motion and the Cunningham correction were neglected [15]. Using the transformations proposed in [15], the following dependence was obtained allowing the calculation of particle fall velocity: 3
C x Re2a ¼
4 da q2g g qa qg 4 ¼ Ar 3 l2g 3 qg
ð8Þ
In the transition regime (1 < Rea < 1000), which is typical for particles >50 lm, the equation Rea ¼ 0:152 Ar0:715 can be applied, whereas, in the laminar regime (Stokes law Cx = 24/Re), Rea = Ar/18.To determine the fall velocity of particles with a 3 diameter da, the Archimedes number Ar ¼ 9:81 da qg ðqa qg Þ=l2g and the corresponding Reynolds number Rea must be calculated:
waf ¼
lg Rea da qg
The results of this calculation are given in Table 6.
ð9Þ
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τ = 30 min
τ = 15 min
τ = 45 min
τ = 75 min
Fig. 6. The test tube bank at the given times (case s).
Table 6 Fall velocity of versus particle diameter. Particle diameter da (lm) Fall velocity wa f (m/s)
300 2.3
210 1.43
147.5 0.95
103.4 0.63
72.5 0.42
50.8 0.3
35.6 0.1
25 0.05
The ash from the container of the test stand (Fig. 4) falls down to the feeder, where it is entrained to the controlled air stream with a starting velocity of wa = 0. In CFD modelling, the streams of air (from the supply fan and the blowing nozzle) as well as the stream and velocity of the solid phase must be given. Given the calculation time, the area of the 3D model was limited to the test duct section (Fig. 4b). A grid of 1.4 million elements was used. To determine the velocity of the particles entering the test duct, a 4-m-long empty vertical duct was modelled. The particles are entrained to the air stream there. Modelling enabled the analysis of the calculation results concerning the final particle velocities, including the influence of gravitational forces. Despite gravity, the influence of the particle shape should be assumed in the model. Fluent allows for the analysis of spherical particles, non-spherical particles and very fine particles smaller than 10 lm if the Cunningham correction is considered. The following data were introduced: air stream corresponding to its actual velocity at the inlet to the test duct win = 10.8 m/s, polydispersion index n = 0.912, ash particles are spherical, minimum particle diameter da min = 25 lm,
Table 7 Measurement results (after 30 min; see Tables 4 and 5) and CFD modelling results at y = 650 mm (Dmt for the tube (3) in the first row; see Fig. 4b). Blowing nozzle
On Off
Concentration
High Low Low
Measurement
Fluent – traverse y = 650 mm
Mass loss (g) Dm t
Velocity max (m/s) wmax
Velocity mean (m/s) w
Conc. max (kg/m3) cmax
Conc. mean (kg/m3) c
Subscripts – i/j
0.3815 0.2292 0.002124
23.1 24.3 5.89
6.61 7.10 5.10
1.4 0.712 0.677
0.442 0.247 0.248
1/1 ½ 2/2
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maximum particle diameter da max = 300 lm, the solid phase has eight characteristic sizes depending on n, da min and da max (25, 35.6, 50.8, 72.5, 103.4, 147.5, 210.4, 300) lm, initial particle velocity wa = 0, the gas phase is air regarded as an ideal gas. The Fluent code determines the eight characteristic diameters from the Rosin–Rammler equation. At each grid element at the inlet, eight characteristic particles are entrained to the air. The final velocities of these particles are presented in Fig. 7. The diagram shows the fall velocities of particles (the final velocities of particles were subtracted from the air velocity at the inlet to the test duct, win = 10.8 m/s). The fall velocities are very close to the values presented in Table 6, which confirms the good quality of the assumptions used for CFD modelling. Fig. 7 shows that, after a trajectory of only 2 m, the ash particles with diameters between 25 lm and 147.5 lm move with velocities similar to the inlet air velocity. A few larger particles reach this velocity after approximately 4 m of flow [14]. In conclusion, the individual particle velocities at the inlet cross-section of the test stand determined with the use of Fluent code were properly estimated and, in practice, were equal to the air velocity. After the preliminary calculations described above, the CFD modelling of the test duct shown in Fig. 4b was performed. The Navier–Stokes (RANS) equations were solved using Fluent. Three turbulence models, which may be used for various CFD applications [16–19], were compared: Spalart-Allmaras (SA) one equation model k–e standard two equation model k–x (Shear–Stress–Transport, SST) two equation model. For the models analysed, the following conditions were adopted:
air and solid phase streams at the inlet of test duct – according to Table 3, stagnation temperature t0, polydispersion index n = 0.912, particles are spherical, particle diameters: da min = 25 lm, da max = 300 lm, the solid phase has eight characteristic sizes as above, initial particle velocity is equal to the air velocity, the gas phase is air regarded as an ideal gas.
Significant mass losses were noted for the central tube of the 1st tube row (1C), as shown in Fig. 4b, due to the operation of the air nozzle. Thus, the analysis was performed for only the first row of the tube bank tested at the level y = 650 mm. The maximum velocity measured in the compressed air nozzle was approximately 38 m/s at the level y = 500 mm (Fig. 4b). Only the SA-model gives this value, as shown in Fig. 8. The other models produced higher values: in the k–x model,
Fig. 7. Final velocities of particles wa = f(da).
W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
483
Fig. 8. Gas-phase velocity distribution at various traverses of the test duct (y = 500, 650, 750 mm).
the maximum velocity was approximately 45 m/s while, in the k–e model, it was greater than 50 m/s. According to the results of the comparison performed, the SA-model was chosen for further analysis. Fig. 9 presents the distribution of the static pressure and velocity along the test duct between the nozzle outlet and y = 650 mm. The mass of the entrained ash was evenly distributed on each grid element of the inlet cross section. This section consists of 464 elements, and 8 particles with various diameters (between 25 and 300 lm) are transported. The erosion rate depends on the velocity, the concentration, the particle diameter, and the element of the inlet section where the ash is entrained to the air. Fig. 10 depicts example trajectories of particles in an element of the first row of the test tube bank when the blowing nozzle is switched on and off. Table 7 shows the data necessary to analyse eq. (10). The values with the subscript ‘max’ are related to the central tube (3) of the first row, as shown in Fig. 4b, because the most intense velocity gradients are in this area. The remaining values in Table 7 are the mean values in the cross-section at y = 650 mm. The transformation of Eq. (5) gives
dbl ¼
n bl cbl bw bl n bcbl Dhbl w ¼ c Dh bw bc w
ð10Þ
The following equations introduce subscripts (i/j) denoting the following cases: i = 1 – blowing nozzle on, i = 2 – blowing nozzle off, j = 1 – high concentration of solid phase, j = 2 – low concentration of solid phase. Consequently, we describe the following combinations: i/j = 1/1 – blowing nozzle on and high concentration of solid phase, i/j = 1/2 – blowing nozzle on and lower concentration of solid phase, i/j = 2/2 – blowing nozzle off and lower concentration of solid phase. Next, the following cases were considered:
0w 1n c max1=1 max1=1 1=1 n c1=1 1=1 Dh1=1 w w @w A c c1=1 ¼ max2=2 max2=2 2=2 c2=2 w Dh2=2 c w
ð11Þ
0w 1n c max1=2 max1=2 1=2 n c1=2 1=2 Dh1=2 w w @w A c c1=2 ¼ max2=2 max2=2 2=2 c2=2 Dh2=2 w c w
ð12Þ
2=2
2=2
2=2
2=2
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Fig. 9. The distribution of static pressure and velocity along the test duct between the nozzle inlet and y = 650 mm.
blowing nozzle on
blowing nozzle off
Fig. 10. Trajectories of an element in the traverse of the first row of tubes.
W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
0w 1n c max1=1 max1=1 1=1 n c1=1 1=1 Dh1=1 w w @w A c c1=1 ¼ max1=2 max1=2 1=2 c1=2 Dh1=2 w c w 1=2
485
ð13Þ
1=2
Eqs. (11)–(13) were applied, and the value of the exponent n was calculated. According to the data given in Table 7:
dbl ¼
Dh1=1 ¼ 179:6 n ¼ 3:27 Dh2=2
dbl ¼
Dh1=2 ¼ 107:9 n ¼ 3:27 Dh2=2
dbl ¼
Dh1=1 ¼ 1:664 n ¼ 3:29 Dh1=2
All results are equal to approximately 3.3. Similar values are obtained in measurements [1–3]. 4. The influence of blowing pressure on velocity distribution Based on the mass loss measurements shown in [14] and the numerical calculations, it can be confirmed that the gas velocity has the greatest impact on particulate erosion. This velocity depends on the stagnation parameters (t0, p0) at the inlet of the nozzle. In Fig. 9, the test stand conditions simulated with use of the CFD model are presented. The inlet pressure was in the range of 3 bar, which corresponds to the values measured in the test stand and given in Table 2. The pressure was calculated using the continuity and energy conservation equations by assuming the air stream, the stagnation temperature t0 and the initial static overpressure. Next, the velocity distribution along the axis of the nozzle stream was calculated. In addition to the test stand value of 3 bar, the calculations were performed for higher (6.3; 12.5 and 20 bar) and lower (1.5 bar) inlet pressures; see Fig. 11 [14]. The diagrams show that the higher the nozzle inlet pressure, the higher the nozzle outlet pressure. Therefore, as a result of the secondary expansion, the velocity of the gas entering the tube bank is also higher. Rapid gas retardation can be seen here. Nevertheless, the gas velocity is still higher than in the case of air nozzle being switched off. The section between the nozzle outlet (y = 320 mm in Fig. 4b) and the point immediately in front of the third tube in the first row of the tube bank (y = 650 mm) was analysed. The length of this section is 330 mm, i.e., the actual distance between the sootblower nozzle and the tube bank in the boiler. The final velocity at y = 650 mm was measured at 24.3 m/s (Table 7 – low concentration), while the pressure was 3 bar. With 6.3 bar, this velocity was approximately 40 m/s; with 12.5 bar, it was 70 m/s; and at 20 bar, it can exceed
Fig. 11. Velocity distribution at the nozzle outlet depending on pressure.
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
Fig. 12. End part of the sootblower RKS 81 E [20].
84 m/s. In all cases except that of 1.5 bar, for pam/p < bL = 0.528, the velocity in the nozzle axis at the outlet amounts to approximately 300 m/s and corresponds to the velocity of sound in the air (Fig. 11). The exponent n for paint erosion according to Eqs. (11)–(13) does not differ significantly from the exponent noted for steel erosion [14]. Therefore, the experimental results may be related to the process occurring in the real boiler of the Opole = 5.2 m/s is the same as that in the test stand (blowing nozpower plant under the assumption that the gas phase velocity w zle switched off). For the typical coal used in the power station, the mean concentration of fly ash in the flue gas is in the range of 15 g/m3. Near the banks of the steam superheaters, RKS 81E blowers manufactured by the Clyde Bergemann company are used in the BP 1150 boiler [20–22]. The rate of the lance travel, extending and retracting (1 blowing cycle), is sc bl = 628 s. The blower takes in approximately 800 kg of steam per cycle. The blower lance makes a rotary motion, during which steam jets from 3 nozzles blow away the ash deposits on the tube banks. The inside diameter of the blower lance is 80.9 mm. Each blower has 3 nozzles with a diameter of approximately 21.1 mm (Fig. 12), which gives a total steam outflow area of FL = 0.00105 m2. The steam pressure is approximately in the range of (0.6–1.0) MPa [21]. This type of pressure range involves a very high velocity of steam at the outlet of the nozzles. The steam jet blower flow characteristics: 800 _ bl ¼ smst ¼ 628 steam mass flow m ¼ 1:274 (kg/s) c bl
lance section area F l ¼
pd2l 4
2
¼ p0:0809 ¼ 0:00515 ðm2 Þ 4
superheated steam parameters at the blower inlet
p ¼ 1 MPa i ¼ 3264 ðkJ=kgÞ; q ¼ 3:26 ðkg=m3 Þ; s ¼ 7:465 ðkJ=kg KÞ t ¼ 400 C _
bl steam flow velocity in the blower lance w ¼ m ¼ 75 (m/s) Fq l
stagnation parameters of superheated steam for isentropic change s = idem
i0 ¼ i þ
w2 ¼ 3267 ðkJ=kgÞ ) 2
t 0 ¼ 401:5 C;p0 ¼ 1:01 MPa
.The equation of continuity implies that the maximum value, (wq) = max, determined by means of a trial method and subject to the condition of constant entropy s0 = sL, occurs at the smallest cross-section of the nozzle. The calculation results are listed in Table 8. The steam velocity at the nozzle outlet is calculated according to the following dependence [23]: Table 8 Characteristic parameters of the steam at the exit of the blowing nozzle. pL bar
iL kJ/kg
kg/m3
qL
wL m/s
wLqL kg/m2 s
9 8 7 6 5.4 5
3232 3197 3159 3116 3088 3067
3.012 2.740 2.475 2.198 2.024 1.908
265 373 465 549 598 632
798 1022 1151 1207 1212 1206
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489 Table 9 and bc ; bw . . Tube thickness loss versus mean flue gas velocity w Sootblower on bc = 1.2 bw = var Air p=3 Bar bl ðm=sÞ w 5.2 ðm=sÞ w 5.2
wL ¼
wmax ðm=sÞ Dh (mm)
Dh (mm)
Superheated steam p = 6.3
p = 10.5
24 40 70 0.1887 0.874 4.71 Sootblower off bc = 1.2; bw = 1.25 ðwmax ¼ 1:25 wÞ 0.00374
p = 20
p = 6 bar
84 8.14
100 13.6
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2ði0 iL Þ
ð14Þ
The superheated steam leaves the nozzle with a velocity of wL 600 m/s at a pressure of pL = 5.4 bar when p0 = 10 bar (Table 8) pL = 3.3 bar when p0 = 6 bar. The difference between the pressure forces in the smallest section and the ambient pressure forces results in an increase of the stream momentum and, consequently, in an increase of the jet velocity outside the nozzle to the value w2.
w2 ¼ wL þ
FL ðp pam Þ105 _ bl L m
ð15Þ
If the pressure at the nozzle inlet is approximately 10 bar, the velocity resulting from expansion can reach values of up to 960 m/s (and w2 780 m/s at 6 bar). However, the jet of steam or air is slowed down rather abruptly at a short distance. The process is presented in Fig. 10 in relation to the air (perfect gas) flow for different values of the pressure [14]. Even after being slowed down by the start of the nozzles (the blower operation), the jet velocities are definitely higher than the velocity of the flow of dust-laden flue gases through the bank with shut down nozzles. The charts show that the higher the pressure before the nozzle, the higher the pressure after it. Thus, due to secondary expansion, the velocity of the gas agent in the area before the tube bank is also higher. The outermost tubes of the neighbouring stages of the steam superheaters, between which the steam jet ash blower is installed, are approximately 2 330 mm. Fig. 11 shows that the air jet velocity for y = 650 mm at a pressure of 10 bar is approximately 60 m/s, with a maximum velocity after secondary expansion of w2 = 510 m/s. In the case of the steam jet blower, where the velocity is w2 = 960 m/s after secondary expansion, the steam velocities before the tubes reach yet higher values. Therefore, for the analysis of the erosion caused by the temporary operation of the steam jet blower, the maximum values of the blowing agent are assumed to be at the level of wmax = 100 m/s. The erosion rate ve bl related to the use of blowers is unusually high, mainly due to the huge velocities given to the ash particles by the blowing agent. The factor that reduces the material loss is the relatively short annual time sbl a of the duration of the process described above compared to the annual operating time of the boiler. Below, a calculation of time sbl a is performed for the conditions prevailing in the area of the steam superheater of the BP1150 boiler. Due to the different dimensions of the test stand and the real boiler duct, the ratio of sootblowing stream in the total gas phase stream, which, in real conditions, consists of flue gas and (while sootblowing) steam, is much smaller for the actual at the level y = 650, as shown in Table 7, occurred boiler. Therefore, noticeable changes in the total gas phase velocity w significantly. Thus, in the test stand, while in the boiler, switching the nozzle on and off does not change the mean velocity w is lower for the test stand in contrast to the corresponding the parameter describing the velocity distribution bw ¼ wmax =w parameter calculated for the actual boiler and the operating sootblower. Nevertheless, the maximum velocity for both cases will only depend upon the operating parameters of the sootblower (air nozzle). According to [10], an abnormal boiler operation, i.e., without sootblowing, was assumed with factors of bw = 1.25 and bc = 1.20–1.37. From the start of the blowing cycle, bw changes dramatically because of the extremely high velocities of the cleaning medium, as shown in Fig. 11. Table 9 shows the tube thickness loss Dh (lm) according to Eq. (1) for a mean = 5.2 m/s, bc = 1.20 and bw = 1.25 at various blowing velocities bw. duct velocity of w Fig. 13 depicts the range of thickness loss for two extreme sets of flue gas parameters when the sootblower is not oper = 5.2 m/s, bw = var, bc = 1.20 and w = 10 m/s, bw = var i bc = 1.37. ating: w As shown in Table 9, the sootblower operation increases the thickness loss (Dh) by approximately 50 times at wmax = 24 m/s and by approximately 3640 times at wmax = 100 m/s compared with normal boiler operation without sootblowing. This analysis shows that the pressure of the blowing steam, which influences its velocity, has a strong effect on the erosion rate, in particular when the distance between the sootblower nozzle and the tube bank is small. Thus, the optimisation of the steam pressure is crucial for the proper operation of sootblowing systems. If the blower is started three times per 24 h, which is considered the highest frequency used in practice, within one year (which, for the Polish power plants, corresponds to approximately sa ffi 6000 h), the annual time of exposure is sbl a = 8.5 h [14].
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
¼ 10 m=s, bw = var and bc = 1.37 - - - - w ¼ 5:2 m=s, bw = var and bc = 1.20. Fig. 13. Tube thickness loss versus maximum velocity wmax —— w Table 10 and bc, bw. Tube thickness loss versus mean flue gas velocity w Sootblower on – sbl a = 8.5 h/a bc = 1.2 bw = var Air p = 3 bar bl ðm=sÞ w 5.2 ðm=sÞ w 5.2
bl ðm=sÞ w 10 wðm=sÞ 10
wmax ðm=sÞ Dh (mm)
Dh (mm) P Dh (mm)
wmax ðm=sÞ Dh (mm)
Dh (mm) P Dh (mm)
Superheated steam p = 6.3 bar
p = 10.5 bar
24 40 70 0.0016 0.0074 0.0396 Sootblower off – sn bl a = 5991.5 h/a bc = 1.2; bw = 1.25 0.0224 0.024 0.03 0.062 Sootblower on – sbl a = 8.5 h/a bc = 1.37 bw = var Air p = 3 bar p = 6.3 bar p = 10.5 bar 24 40 70 0.0021 0.01 0.053 Sootblower off – sn bl a = 5991.5 h/a bc = 1.37; bw = 1.25 0.213 0.215 0.223 0.266
p = 20 bar
p = 6 bar
84 0.069
100 0.116
0.091
0.138
p = 20 bar 84 0.092
Superheated steam p = 6 bar 100 0.155
0.305
0.368
It may be assumed that the boiler banks are exposed to a more intense erosion for 8.5 h of the 6000 h of operating time per year. With these assumptions, the tube material loss according to formula (1) is the total of the loss without blower operation (sn bl a = 5991.5 h) and the loss with blower operation for 8.5 h according to formula (3). Example calculation results are listed in Table 10. The steam expansion behind the nozzle is related to the lowering of the temperature. However, there is no danger of creating wet steam with water droplets, which might erode the heating surfaces of the boiler. In the worst case, assuming isentropic expansion from pL = 5.4 bar to ambient pressure pam = 1 bar, we obtain a final temperature of approximately 120 °C. For pL = 3.3, this temperature amounts to 170 °C. In actual conditions (irreversible process), the final temperature will be even higher. Furthermore, the blowing steam is introduced to the hot flue gas and sustains the steam in a superheated state. 5. Conclusions 1. The laboratory experiments and numerical modelling showed clearly that the mechanism of sootblower erosion consists of entraining the fly ash particles from the flue gas into the high-velocity blowing medium stream. Our calculations indicate that there is no danger of creating wet steam containing water droplets while steam expands and the tube wear as a result of the impact of water droplets flowing with the steam shall be neglected. 2. The conformity (with accuracy sufficient for most engineering purposes) between CFD modelling and experiments proves that the numerical modelling may be used as a reliable tool for a quantitative estimation of sootblower erosion in power boilers. 3. Approximate calculations showed that the sootblower erosion rate is many times higher than the fly ash erosion during a normal boiler operation. Therefore, even relatively short sootblower operation periods may lead to serious damages to the heating surfaces. 4. An important parameter of influence is the pressure of the sootblowing medium. Thus, the lowest pressure ensuring satisfactory cleaning effect and minimum thickness loss should be determined experimentally. References [1] Meuronen V. Ash particle erosion on steam boiler convective section. Lappeenranta University of Technology, Research Papers 64, Lappeenranta; 1997. _ [2] S´wirski J. Studies on ash erosion and the assessment tube erosive wear in boilers (Badania erozji popiołowej i ocena zuzycia rur kotłowych wskutek jej działania). Scientific books of the Institute of Power Engineering, Book 1, Warsaw; 1975 [in Polish].
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[3] S´wirski J. Design guidelines against fly ash erosion for boiler heating surfaces and flue gas ducts (Wytyczne projektowania kanałów spalinowych i powierzchni ogrzewalnych kotłów dla ochrony przed erozja˛ popiołowa˛). Report no. 10940 of the Institute of Power Engineering, Warsaw; 1974 [in Polish]. [4] Uetz H. Strahlverschleis, Mitt. der VGB 49; 1969. [5] Pronobis M. Operational measures to reduce erosion in convection surfaces of boilers (Przedsie˛wzie˛cia eksploatacyjne dla zmniejszenia erozji kotłowych pe˛czków konwekcyjnych). Scientific books of the Silesian University of Technology, Book no. 121, Gliwice; 1994 [in Polish]. [6] Urban B. Effect of tube bundle arrangement on fly ash erosion of heating surfaces (Vliv usporˇádáni trubek ve svazku vy´hrˇevne plochy na jejich popilkovy´ otr). Energetika no. 3; 1980 [in Czech]. [7] Fehndrich W. Verschleißuntersuchungen an Kesselrohren, Mitt. der VGB 49; 1969. [8] Pronobis M, Cˇech B. Effect of local flue gas parameters on erosion of convection heating surfaces (Wpływ lokalnych parametrów spalin na erozje˛ konwekcyjnych powierzchni ogrzewalnych). In: 9th International conference on boiler technology 2002 scientific books of the silesian university of technology no. 10; 2002 [in Polish]. [9] Kadlec Z, Cˇech B, Roubicˇek V, Kolat P. Diagnostic methods for operating surveillance of large fluidized bed boilers. PowerPlant Chem 2007;9(6):381. [10] Pronobis M, Wojnar W. Preliminary calculations of erosion wear resulting from exfoliation of iron oxides in austenitic superheaters. Eng Fail Anal 2013;32:54–62. [11] Wejkowski R, Pronobis M. Investigations of ash erosion intensity in selected convective surfaces of boilers (Badania intensywnos´ci erozji popiołowej w wybranych powierzchniach konwekcyjnych kotłów). Chem Process Eng 2004;24 [in Polish]. [12] Fluent. Computational fluid dynamics, FLUENT Inc., England. [13] Parslow GI, Stephenson DJ, Strutt JE, Tetlow S. Paint layer erosion resistance behavior for use in a multilayer paint erosion indication technique. Wear 1997;212. _ ´ termicznych i erozyjnych [14] Wojnar W. Identification of thermal and erosion hazards to reduce superheater failures (Identyfikacja zagrozen przegrzewaczy pary w celu zmniejszenia ich awaryjnos´ci). PhD thesis. Institute of Power Engineering and Turbomachinery, Silesian University of Technology, Gliwice; 2008 [in Polish]. [15] Kabsch P. Dust extraction and dust collectors. Volume 1. Mechanics of aerosols and dry dust collectors (Odpylanie i odpylacze. Tom 1. Mechanika aerozoli i odpylacze suche). WNT, Warsaw; 1992 [in Polish]. [16] Tandra D, Kaliazine A, Cormack DE. Honghi tran: numerical modeling of sootblower jet flow in a kraft recovery boiler. Canadian pulp and paper graduate students seminars; 2004 [Internet]. [17] Sciubba E, Zeoli N. A study of sootblower erosion in waste – incinerating heat boilers. Universita di Roma, Journal of Energy Resources Technology, 2006. [18] Xiao Q, Tsai HM, Papamoschou D. Numerical study of jet plume instability from an overexpanded nozzle. AIAA 2007-1319; 2007. [19] Tandra D, Kaliazine A, Cormack DE. Honghi tran: numerical simulation of supersonic jet flow using a modified k-model. Int J Comput Fluid Dynam 2006;20(1):19–27. [20] http://www.clydebergemannpowergroup.com. [21] Pietrzyk M. Optimization of sootblowing of BP-1150 boiler heating surfaces (Optymalizacja zdmuchiwania powierzchni grzewczych kotła BP-1150). In: 10th Power engineering seminar, scientific books of the silesian university of technology no. 7; 2001 [in Polish]. [22] Jameel M. Steam saving in recovery boilers. Bergemann USA. [23] Szargut J. The theory of thermal processes (Teoria procesów cieplnych). PWN. Warsaw; 1973 [in Polish].
Contents lists available at SciVerse ScienceDirect
Engineering Failure Analysis journal homepage: www.elsevier.com/locate/engfailanal
Erosion of heat exchangers due to sootblowing Wacław Wojnar ⇑ Silesian University of Technology, Faculty of Energy and Environmental Engineering, Institute of Power Engineering and Turbomachinery, Division of Boilers and Steam Generators, Konarskiego 20, 44-100 Gliwice, Poland
a r t i c l e
i n f o
Article history: Received 20 May 2013 Received in revised form 21 June 2013 Accepted 22 June 2013 Available online 4 July 2013 Keywords: Erosion Superheaters Sootblower CFD
a b s t r a c t The erosion rate of convection surfaces is crucial for boiler reliability. This paper presents the impact of the action of a sootblowing stream (steam and air) jet on the sustainability of tubes. The model investigations described here compared the erosion rates of a tube bank while turning on or off an additional jet stream of compressed air that simulates the operation of a sootblower. The numerical model code Fluent was used. Based on mass loss measurements and numerical calculations, it can be confirmed that the gas velocity has the greatest impact on the particulate erosion. This velocity depends on the stagnation parameters at the inlet of the nozzle. This paper presents calculations for the erosion loss of a heat exchanger built in the actual object. Ó 2013 Elsevier Ltd. All rights reserved.
1. Introduction In boilers that fire solid fuels, various erosion processes can occur, leading to serious operational and economic problems. The wear of the heat exchanger tubes depends, among other factors, on: The velocity of solid particles, which is related to the erosion rate as a power function. The experimentally determined powers have been compared [1], and they amount to 2.4–3.5 for the wear of metals, approximately 3 for ceramics and approximately 5 for polymers. The concentration of the solid phase in the gas. At low concentrations, the erosion rate increases with the concentration. When the concentrations are high, the influence of this parameter may be insignificant or even decrease the wear observed due to collisions between the incident and reflected particles. The particle diameter: the erosion rate increases until the particle diameter reaches the range of 50–100 lm. Above this range, the erosion rate is not influenced by the particle diameter. The shape of particles: for metallic surfaces, non-spherical, sharply edged particles cause stronger wear. The hardness of the erodent. The impact angle ue between the two-phase stream and the surface: the maximum wear of ductile materials occurs for ue max = 20°–40° [2,3], while that of brittle materials occurs for perpendicular impact. The erosion rates of the convection surfaces are crucial for the boiler reliability. Many papers, for example [1–8], have described tube failures resulting from wear due to the fly ash in the flue gas, but the erosion caused by particulate matter entrained in the sootblowing stream (steam or air) has not been investigated at the same level. Because the existing methods
⇑ Tel.: +48 322372374; fax: +48 32 2372193. E-mail address: [email protected] 1350-6307/$ - see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.engfailanal.2013.06.026
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
Nomenclature Ar Cx c cbl cmax da da min da max dbl dl D ee Fl FL Fr Fu g i i0 iL ma m_ a _ bl m mst n p P p0 p1 pam pc pL Re Rea s s1 s2 t0 y z1 z2 ve = Dh/s ve bl = Dhbl/sbl-a V V_ V_ n w w2 wa waf wg win wL wm wmax w wbl
Archimedes number współczynnik oporów ruchu ziarna pyłu w gazie, mean concentration of fly ash in the flue gas (sootblower off), g/m3 mean concentration of fly ash in the flue gas (sootblower on), g/m3 maximum concentration of fly ash in the flue gas, g/m3 particle diameter, lm minimum particle diameter, lm maximum particle diameter, lm blower nozzle diameter, m inside diameter of the blower lance, m outer tube diameter, m erosion coefficient defining the erosive properties of the fly ash, lm/h blower lance cross-section, m2 total steam outflow area, m2 resistance force of the spherical particle with diameter da, N uplift force of the gaseous medium with the density qg, N acceleration of gravity, m/s2 enthalpy, kJ/kg stagnation enthalpy, kJ/kg steam enthalpy at the blower nozzle outlet, kJ/kg solid phase mass (erodent mass), kg, erodent stream, kg/s steam mass flow from the blower, kg/s steam mass per 1 blowing cycle, kg polydispersion index steam pressure at the blower inlet, bar gravity force of solid phase mass ma, N stagnation pressure, Pa pressure on the inlet of the nozzle, bar, ambient pressure, bar, pressure on the outlet of the compressor, bar steam pressure at the blower nozzle outlet, bar Reynolds number, Reynolds number for particle fall entropy, kJ/kgK transverse pitch, mm longitudinal pitch, mm stagnation temperature, °C height coordinate, mm number of tubes in the first row number of rows, erosion rate, lm/h blower operation-related erosion rate, lm/h grid volume, m3 supply fan output, kg/s, m3/s air stream in the nozzle, kg/s steam flow velocity in the blower lance, m/s blowing agent velocity outside the nozzle, m/s particle velocity, m/s fall velocity, m/s flue gas velocity in the free section at the first tube, m/s inlet velocity of the test duct (level y = 700 mm, Fig. 4b), m/s steam velocity at the blower nozzle outlet, m/s velocity in the cross-section of the nozzle (Fig. 2), m/s maximum velocity of the dust-laden flue gases in the duct section, m/s mean flue gas velocity in the empty duct at the bank inlet (sootblower off), m/s mean flue gas velocity in the empty duct at the bank inlet (sootblower on), m/s
W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
475
Greek symbols:
amax b2 bc bL bm bw dbl Dh Dhbl Dmtbl Dmt Dp
gp lg q qa qg qL qt
r1 ¼ s1 =D r2 ¼ s2 =D s sn bl a sc bl sa sbla u ue ue max
angle spanning over eroded tube perimeter parameter defining the impact of longitudinal pitch of the bank on the erosive wear of the tubes parameter defining the fly ash concentration distribution in the flue gases critical pressure ratio parameter defining the impact of the operating temperature and the steel grade on the erosive wear parameter defining the flue gas velocity distribution in the boiler duct factor introducing the sootblowing effect thickness loss of the tube, lm, mm thickness loss of the tube (sootblower on), lm, mm test tube mass loss at air jet switched on (g) test tube mass loss at air jet switched off (g) differential pressure – orifice plate, Pa mean probability of ash particle impingement on the tube gas dynamic viscosity, kg/m s density, kg/m3 erodent density, kg/m3 gas density, kg/m3 steam density at the blower nozzle outlet, kg/m3 tube density, kg/m3 relative transverse pitch of the tube, m relative longitudinal pitch of the tube, m time interval, s annual time with the blower shut down, h blowing cycle duration time, s annual boiler operating time, h annual duration time of the erosion process caused by ash blowing, h air humidity, % angle between two-phase stream and the surface tube, ° maximum angle between two-phase stream and the surface tube, °
of calculation are not appropriate for this case, this work is an attempt to determine this phenomenon quantitatively to reduce the erosive damage (Fig. 1) by controlling the process parameters.
2. Calculation of the erosion rate during sootblowing The fly ash erosion rate (lm/h) of the plain tube banks in a flue gas stream may be described, according to [2,9], as
3 b3w ee cbc bm b2 gp Dh=s ¼ v e ¼ 2:33 107 w
ðlm=hÞ
cmax where bw ¼ wmax ; bc ¼ w c The method to calculate the parameters in Eq. (1) is presented in [10].
Fig. 1. Examples of erosion damage to convection surfaces.
ð1Þ
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
(a)
(b)
Tube bank
w
Flue gas with ash particles
wbl
wmax Sootblower nozzle
w max Sootblower nozzle
wm
¼ 5—10 m=s;wmax 1:25 wÞ and (b) the blowing jest turned on (w bl ffi w, Fig. 2. Scheme for the erosion rate calculation: (a) the blowing jest turned off ðw wmax = f(static parameters of the blowing medium p0, t0; distance nozzle-tubes; wm), wm – velocity in the cross-section of the nozzle, which is ca. 300 m/s for air and ca. 600 m/s for steam).
The problem investigated in this work differs from typical particle erosion problems on convective sections of steam boilers. An additional element is the jet stream of a high velocity soot-blowing medium, which entrains and accelerates the ash particles, directing them towards the tubes. The model investigations described below compare the erosion rates of a tube bank while turning the additional jet stream of the compressed air on or off, thereby simulating the sootblower operation. Assuming equal periods of rig operation in both states, the factor describing the sootblower effect is introduced as
dbl ¼
Dmtbl Dmt s¼idem
ð2Þ
Because the basic formula (1) describes the erosion loss as a reduction of the tube wall thickness in the area of the most intense erosion, the same formula can be used to determine sootblowing-induced erosion (3) for the case when the sootblowing nozzle is turned on. Both cases are presented in Fig. 2.
Dhbl
sbl a
3 b3w ee cbc bm b2 gp Þ ¼ v ebl ¼ 2:33 107 ðw bl
ðlm=hÞ
ð3Þ
As shown in Fig. 6, during sootblowing, the area of the most intense wear on the tubes occurs at an angle of approximately 40° symmetric about the inflow axis. Many investigations have shown that the maximum wear at this location also occurs during typical fly ash erosion without blowing. Assuming that the thickness loss Dh occurs along the radius R (from R1 to R2)
Fig. 3. Indication of tube wall thickness losses.
W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
(a)
477
(b)
Fig. 4. (a) Layout of the test stand: 1 – tube bank, 2 – test duct, 3 – fan, 4 – erodent container, 5 – feeder, 6 – cyclone precipitator, 7 – compressor, 8 – orifice plate, 9 – rubber tube, 10 – manometer, 11 – nozzle, 12 – orifice plate and (b) test duct with traverses.
and that the area of the most intense wear is located at the section of the perimeter described by the angle amax (Fig. 3), the following equation can be used:
dbl ¼
Dmt bl DV tbl qt amax ðR21 R22 Þtbl Dhbl ð2R1 Dhbl Þ ¼ ¼ ¼ Dmt s¼idem DV t q t Dhð2R1 DhÞ amax ðR21 R22 Þt
ð4Þ
while R1 Dh, this equation becomes
dbl ¼
Dmt bl Dhbl Dmt s¼idem Dh
Eventually, it can be assumed that the factor dbl
dbl ¼
Dhbl Dmt bl ¼ Dh s¼idem Dmt s¼idem
ð5Þ
proves that the mass loss measurements is equal to erosion rate and reflects the tube wall thickness loss for a given time period.
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
2.1. Test rig and the method of investigation At the Institute of Power Engineering and Turbomachinery of the Silesian University of Technology, there is a test stand for investigating the phenomena of erosion, which is shown in Fig. 4 [1,2,11]. The test stand consists of a fly erodent container (4), a feeder (5), a duct connected to a supply fan (3), a test duct with the tube bank under investigation (1), a cyclone precipitator that (6) recirculates the erodent to the container and an output duct with a stack. The compressed air from two compressors is directed to the test duct (1) below the tube bank using a nozzle (11) with a diameter of dbl = 4 mm located exactly on the axis of the test duct (2). The air stream was measured with an ISA orifice plate (8). The differential pressure on the orifice plate was measured with a U-tube gauge. The air temperatures were measured using mercury thermometers. A pressure gauge (10) measuring the static pressure p1 at its inlet was fixed in front of the nozzle. The measured values of p1 were later compared with those calculated by Fluent [12]. The air stream at the outlet of the supply fan was also measured with an orifice plate (12), as shown in Fig. 4a. The supply air stream was tuned using the control flap located on the fan inlet to reach values in the range of 6–10 m/s at the cross-section of the first row of the tube bank, typical for coal fired boilers. The model steel tubes were covered with blue (outer layer), yellow and red (inner layer) paint according to the idea in [13]. In Fig. 4b, the following traverses (levels according to height – y coordinate) are shown:
y = 320 mm y = 500 mm y = 650 mm y = 715 mm
– – – –
compressed air nozzle outlet, cross-section where the velocity distribution was measured and modelled by Fluent, on this level, using Fluent, the velocities and particle concentrations were estimated to analyse eq. (1), cross-section of the first row of the tube bank.
The test tube bank arrangement was as follows:
external diameter of tubes D = 21.3 mm transverse pitch s1 = 67 mm longitudinal pitch s2 = 90 mm number of tubes in the first row z1 = 5 number of rows z2 = 8.
The air-erodent mixture circulates within this system. The fly ash from a PC boiler, which is described in Table 1, was used as the solid phase. Nearly 80% of the particles remained between 40 and 120 lm, with a polydispersion index of n = 0.912. It was decided that the similarity of the velocity, rather than the similarity of the Re number ðRe ¼ wg Dqg =lg Þ, is preserved in the test rig. In an actual boiler, Re = 2000 – 3000, and this value depends on the flue gas temperature and the velocity, which is in the range of wg = (6–9) m/s in PC boilers. The assumption of preserving the same Re value in the test stand kept at 0 °C, yields velocities equal to approximately 1 m/s, which makes pneumatic transport of the fly ash impossible. Consequently, the test rig is not capable of modelling all similarity criteria in parallel. Nevertheless, it is the velocity that is the most important parameter that influences the particulate erosion. Therefore, in the model duct, the velocities were kept exactly the same, as in an actual boiler. 2.1.1. Discussion of test results Three cases of modelling were analysed [14]: r s t
– – – – – –
air blowing nozzle ‘on’ _ a ¼ 0:06076 kg/s (high concentration) erodent stream m air blowing nozzle ‘on’ _ a ¼ 0:03383 kg=s (low concentration) erodent stream m air blowing nozzle ‘off’ _ a ¼ 0:03383 kg=s: erodent stream m
The operating data for the test stand required to estimate the representative compressed air stream are given in Table 2. Table 1 Characteristics of the erodent (fly ash). Density
Chemical composition (%)
qa
CaO 2.88
2700 kg/m3
MgO 2.41
Fe2O3 12.99
Na2O 0.51
Fineness Rx (%) K2O 1.74
SiO2 60.00
Al2O3 18.89
TiO2 0.57
R0.025 58.8
R0.032 52.2
R0.040 48.5
R0.063 36
R0.088 22
R0.12 11.3
R0.20 2.6
R0.30 0.6
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489 Table 2 Measurement results of the compressed air stream at the nozzle (dbl = 4 mm).
Temperature Pressure on the outlet of the compressor Pressure on the inlet of the nozzle Differential pressure – orifice plate Air stream [14]
Symbol
Unit
Value
t pc p1 Dp V_n
°C bar bar Pa kg/s
8.0 3.7 3.0 2100 0.0092
Table 3 Test stand investigation results. Ambient parameters
r s t
Supply fan output
t
pb
u
°C
mm Hg
%
V_ kg/s
V_ m3/s
10 10 23
752 752 742
55 55 68
0.269 0.269 0.258
0.216 0.216 0.216
Air stream in the nozzle V_n
Stream of the erodent m_ a
Mean concentration c
kg/s
Max. velocity in the axis of the nozzle (y = 500 mm) w500 max m/s
kg/s
g/m3
0.0092 0.0092 0
38 38 5,5
0.06076 0.03383 0.03383
276 154 157
In Table 3, the results concerning the three cases given above are presented, whereas Fig. 4 depicts the velocity distribution in the test duct at the level y = 500 mm, where the velocity probe (WTI) was introduced. Table 3 also shows the ash concentration for a cubic meter of the transported air. According to the investigation results presented in Fig. 5, with the compressed air nozzle switched off, the mean velocity = 5.2 m/s (y = 500 mm), which corresponds to a velocity in the range of wg = 7 m/s in the free section at the in the duct was w first tube row level (y = 715 mm). This velocity range is similar to that in actual boilers. The inlet velocity to the test duct (level y = 700 mm, Fig. 4b) was win = 10.8 m/s. On purpose the first measurement was conducted with the air blowing nozzle switched on and at a high erodent concentration and served as a basis for further comparisons. The purpose was to determine the shortest period after which the lowest (red) paint layer was exposed (the metallic tube surface was still hidden). The time noted was 30 min, and therefore, this period of time was chosen as a reference for further investigation. Exemplary mass losses of tubes after 30 min of exposure to erosion are presented in Tables 4 and 5. With the exception of the mass loss of the central tube, which is shown in Table 4, the mass losses of the tubes were very small. The majority of the results were within the range of the measurement errors, and thus, some losses in Table 4 are negative (mass increase). Therefore, only the central tube was chosen for further investigation. _ a ¼ 0:03383 kg=s). Fig. 6 presents the test tube bank at certain times (air blowing nozzle on and stream of the erodent m During the measurement with the air blowing nozzle switched off, which simulates the boiler duct when the sootblowers are not in operation, a noticeable mass loss of the paint covering the tubes (approximately 0.1 g) was recorded only after 23 h of the test run. The loss after 30 min (the reference period of time) was not measureable with satisfactory precision, and thus, it was necessary to assume a linear, time-dependent mass loss function. With this assumption, the mass loss for the central tube in the first row was Dmt = 0.002124 g/30 min). To determine the factor describing the influence of sootblowing on the erosion rate dbl, which is given in Eq. (3), the distribution of the velocity and the concentration of the erodent is necessary. Both values were calculated using Fluent [12].
8
Velocity [m/s]
7 6 5 4 3 2
1
2
3
4
5
Measuring point Fig. 5. Velocity distribution at the cross-section of the test duct (y = 500 mm) air blowing nozzle out – fan output V_ ¼ 0:216 m3 =s.
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
Table 4 _ a ¼ 0:03383 kg=s (low concentration). Mass losses of test tubes for case 2 – air blowing nozzle on and stream of the erodent m Mass loss Dmt (g) after s = 30 min Row/column
A
B
C
D
E
8 7 6 5 4 3 2 1
0.0077 0.0064 0.0055 0.0112 0.0066 0.0029 0.0068 0.0081
0.0042 0.0064 0.0167 0.0098 0.035 0.0196 0.0137 0.0135
0.0018 0.0029 0.0133 0.015 0.0183 0.021 0.0254 0.2292
0.0075 0.0041 0.001 0.0002 0.0052 0.0049 0.0064 0.0042
0.0111 0.0043 0.0066 0.0077 0.0008 0.0072 0.0051 0.009
Table 5 _ a ¼ 0:06076kg=s (high concentration). Mass losses of test tubes for case 1 – air blowing nozzle on and stream of the erodent m Mass loss Dmt (g) after s = 30 min Row/column
A
B
C
D
E
8 7 6 5 4 3 2 1
0.004 0.0084 0.0114 0.0185 0.0144 0.0115 0.0141 0.005
0.0086 0.0165 0.017 0.0207 0.0323 0.0402 0.0441 0.0178
0.0144 0.0231 0.022 0.0256 0.0221 0.036 0.0291 0.3815
0.0142 0.0164 0.0236 0.0186 0.0144 0.0372 0.037 0.0184
0.0041 0.0084 0.0252 0.0081 0.0062 0.0044 0.0116 0.004
3. Numerical modelling of erosion When modelling erosion in Fluent [12], the description of the concentration is based on the residence time of the particulate matter in the unit of the grid volume V:
c ¼
_ as m V
ðkg=m3 Þ
ð6Þ
_a m Transforming Eq. (6) to c ¼ Fw , where F represents the area of the cross-section of the grid element, shows the connection a between the velocity of the particles and the concentration. The velocity of the particles depends on the gravitational force and may be influenced by the flow direction (upwards or downwards). The flow direction consequently increases or decreases the velocity of the particles, which determines the erosion rate. In the test stand, the air flows upwards, and, according to Stokes law, the particles uniformly accelerate until they reach constant terminal velocities at which the resistance force equals the gravity force of the particle minus the uplift force:
P ¼ Fr þ Fu
ð7Þ
where P ¼ ma g – gravity force of solid phase mass ma; F r ¼ 3plg da wa – resistance force of the spherical particle with diama eter da; F u ¼ m qa qg g – uplift force of the gaseous medium with the density qg. Because the particles are of diameters >10 lm, the Brownian motion and the Cunningham correction were neglected [15]. Using the transformations proposed in [15], the following dependence was obtained allowing the calculation of particle fall velocity: 3
C x Re2a ¼
4 da q2g g qa qg 4 ¼ Ar 3 l2g 3 qg
ð8Þ
In the transition regime (1 < Rea < 1000), which is typical for particles >50 lm, the equation Rea ¼ 0:152 Ar0:715 can be applied, whereas, in the laminar regime (Stokes law Cx = 24/Re), Rea = Ar/18.To determine the fall velocity of particles with a 3 diameter da, the Archimedes number Ar ¼ 9:81 da qg ðqa qg Þ=l2g and the corresponding Reynolds number Rea must be calculated:
waf ¼
lg Rea da qg
The results of this calculation are given in Table 6.
ð9Þ
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τ = 30 min
τ = 15 min
τ = 45 min
τ = 75 min
Fig. 6. The test tube bank at the given times (case s).
Table 6 Fall velocity of versus particle diameter. Particle diameter da (lm) Fall velocity wa f (m/s)
300 2.3
210 1.43
147.5 0.95
103.4 0.63
72.5 0.42
50.8 0.3
35.6 0.1
25 0.05
The ash from the container of the test stand (Fig. 4) falls down to the feeder, where it is entrained to the controlled air stream with a starting velocity of wa = 0. In CFD modelling, the streams of air (from the supply fan and the blowing nozzle) as well as the stream and velocity of the solid phase must be given. Given the calculation time, the area of the 3D model was limited to the test duct section (Fig. 4b). A grid of 1.4 million elements was used. To determine the velocity of the particles entering the test duct, a 4-m-long empty vertical duct was modelled. The particles are entrained to the air stream there. Modelling enabled the analysis of the calculation results concerning the final particle velocities, including the influence of gravitational forces. Despite gravity, the influence of the particle shape should be assumed in the model. Fluent allows for the analysis of spherical particles, non-spherical particles and very fine particles smaller than 10 lm if the Cunningham correction is considered. The following data were introduced: air stream corresponding to its actual velocity at the inlet to the test duct win = 10.8 m/s, polydispersion index n = 0.912, ash particles are spherical, minimum particle diameter da min = 25 lm,
Table 7 Measurement results (after 30 min; see Tables 4 and 5) and CFD modelling results at y = 650 mm (Dmt for the tube (3) in the first row; see Fig. 4b). Blowing nozzle
On Off
Concentration
High Low Low
Measurement
Fluent – traverse y = 650 mm
Mass loss (g) Dm t
Velocity max (m/s) wmax
Velocity mean (m/s) w
Conc. max (kg/m3) cmax
Conc. mean (kg/m3) c
Subscripts – i/j
0.3815 0.2292 0.002124
23.1 24.3 5.89
6.61 7.10 5.10
1.4 0.712 0.677
0.442 0.247 0.248
1/1 ½ 2/2
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maximum particle diameter da max = 300 lm, the solid phase has eight characteristic sizes depending on n, da min and da max (25, 35.6, 50.8, 72.5, 103.4, 147.5, 210.4, 300) lm, initial particle velocity wa = 0, the gas phase is air regarded as an ideal gas. The Fluent code determines the eight characteristic diameters from the Rosin–Rammler equation. At each grid element at the inlet, eight characteristic particles are entrained to the air. The final velocities of these particles are presented in Fig. 7. The diagram shows the fall velocities of particles (the final velocities of particles were subtracted from the air velocity at the inlet to the test duct, win = 10.8 m/s). The fall velocities are very close to the values presented in Table 6, which confirms the good quality of the assumptions used for CFD modelling. Fig. 7 shows that, after a trajectory of only 2 m, the ash particles with diameters between 25 lm and 147.5 lm move with velocities similar to the inlet air velocity. A few larger particles reach this velocity after approximately 4 m of flow [14]. In conclusion, the individual particle velocities at the inlet cross-section of the test stand determined with the use of Fluent code were properly estimated and, in practice, were equal to the air velocity. After the preliminary calculations described above, the CFD modelling of the test duct shown in Fig. 4b was performed. The Navier–Stokes (RANS) equations were solved using Fluent. Three turbulence models, which may be used for various CFD applications [16–19], were compared: Spalart-Allmaras (SA) one equation model k–e standard two equation model k–x (Shear–Stress–Transport, SST) two equation model. For the models analysed, the following conditions were adopted:
air and solid phase streams at the inlet of test duct – according to Table 3, stagnation temperature t0, polydispersion index n = 0.912, particles are spherical, particle diameters: da min = 25 lm, da max = 300 lm, the solid phase has eight characteristic sizes as above, initial particle velocity is equal to the air velocity, the gas phase is air regarded as an ideal gas.
Significant mass losses were noted for the central tube of the 1st tube row (1C), as shown in Fig. 4b, due to the operation of the air nozzle. Thus, the analysis was performed for only the first row of the tube bank tested at the level y = 650 mm. The maximum velocity measured in the compressed air nozzle was approximately 38 m/s at the level y = 500 mm (Fig. 4b). Only the SA-model gives this value, as shown in Fig. 8. The other models produced higher values: in the k–x model,
Fig. 7. Final velocities of particles wa = f(da).
W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
483
Fig. 8. Gas-phase velocity distribution at various traverses of the test duct (y = 500, 650, 750 mm).
the maximum velocity was approximately 45 m/s while, in the k–e model, it was greater than 50 m/s. According to the results of the comparison performed, the SA-model was chosen for further analysis. Fig. 9 presents the distribution of the static pressure and velocity along the test duct between the nozzle outlet and y = 650 mm. The mass of the entrained ash was evenly distributed on each grid element of the inlet cross section. This section consists of 464 elements, and 8 particles with various diameters (between 25 and 300 lm) are transported. The erosion rate depends on the velocity, the concentration, the particle diameter, and the element of the inlet section where the ash is entrained to the air. Fig. 10 depicts example trajectories of particles in an element of the first row of the test tube bank when the blowing nozzle is switched on and off. Table 7 shows the data necessary to analyse eq. (10). The values with the subscript ‘max’ are related to the central tube (3) of the first row, as shown in Fig. 4b, because the most intense velocity gradients are in this area. The remaining values in Table 7 are the mean values in the cross-section at y = 650 mm. The transformation of Eq. (5) gives
dbl ¼
n bl cbl bw bl n bcbl Dhbl w ¼ c Dh bw bc w
ð10Þ
The following equations introduce subscripts (i/j) denoting the following cases: i = 1 – blowing nozzle on, i = 2 – blowing nozzle off, j = 1 – high concentration of solid phase, j = 2 – low concentration of solid phase. Consequently, we describe the following combinations: i/j = 1/1 – blowing nozzle on and high concentration of solid phase, i/j = 1/2 – blowing nozzle on and lower concentration of solid phase, i/j = 2/2 – blowing nozzle off and lower concentration of solid phase. Next, the following cases were considered:
0w 1n c max1=1 max1=1 1=1 n c1=1 1=1 Dh1=1 w w @w A c c1=1 ¼ max2=2 max2=2 2=2 c2=2 w Dh2=2 c w
ð11Þ
0w 1n c max1=2 max1=2 1=2 n c1=2 1=2 Dh1=2 w w @w A c c1=2 ¼ max2=2 max2=2 2=2 c2=2 Dh2=2 w c w
ð12Þ
2=2
2=2
2=2
2=2
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
Fig. 9. The distribution of static pressure and velocity along the test duct between the nozzle inlet and y = 650 mm.
blowing nozzle on
blowing nozzle off
Fig. 10. Trajectories of an element in the traverse of the first row of tubes.
W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
0w 1n c max1=1 max1=1 1=1 n c1=1 1=1 Dh1=1 w w @w A c c1=1 ¼ max1=2 max1=2 1=2 c1=2 Dh1=2 w c w 1=2
485
ð13Þ
1=2
Eqs. (11)–(13) were applied, and the value of the exponent n was calculated. According to the data given in Table 7:
dbl ¼
Dh1=1 ¼ 179:6 n ¼ 3:27 Dh2=2
dbl ¼
Dh1=2 ¼ 107:9 n ¼ 3:27 Dh2=2
dbl ¼
Dh1=1 ¼ 1:664 n ¼ 3:29 Dh1=2
All results are equal to approximately 3.3. Similar values are obtained in measurements [1–3]. 4. The influence of blowing pressure on velocity distribution Based on the mass loss measurements shown in [14] and the numerical calculations, it can be confirmed that the gas velocity has the greatest impact on particulate erosion. This velocity depends on the stagnation parameters (t0, p0) at the inlet of the nozzle. In Fig. 9, the test stand conditions simulated with use of the CFD model are presented. The inlet pressure was in the range of 3 bar, which corresponds to the values measured in the test stand and given in Table 2. The pressure was calculated using the continuity and energy conservation equations by assuming the air stream, the stagnation temperature t0 and the initial static overpressure. Next, the velocity distribution along the axis of the nozzle stream was calculated. In addition to the test stand value of 3 bar, the calculations were performed for higher (6.3; 12.5 and 20 bar) and lower (1.5 bar) inlet pressures; see Fig. 11 [14]. The diagrams show that the higher the nozzle inlet pressure, the higher the nozzle outlet pressure. Therefore, as a result of the secondary expansion, the velocity of the gas entering the tube bank is also higher. Rapid gas retardation can be seen here. Nevertheless, the gas velocity is still higher than in the case of air nozzle being switched off. The section between the nozzle outlet (y = 320 mm in Fig. 4b) and the point immediately in front of the third tube in the first row of the tube bank (y = 650 mm) was analysed. The length of this section is 330 mm, i.e., the actual distance between the sootblower nozzle and the tube bank in the boiler. The final velocity at y = 650 mm was measured at 24.3 m/s (Table 7 – low concentration), while the pressure was 3 bar. With 6.3 bar, this velocity was approximately 40 m/s; with 12.5 bar, it was 70 m/s; and at 20 bar, it can exceed
Fig. 11. Velocity distribution at the nozzle outlet depending on pressure.
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
Fig. 12. End part of the sootblower RKS 81 E [20].
84 m/s. In all cases except that of 1.5 bar, for pam/p < bL = 0.528, the velocity in the nozzle axis at the outlet amounts to approximately 300 m/s and corresponds to the velocity of sound in the air (Fig. 11). The exponent n for paint erosion according to Eqs. (11)–(13) does not differ significantly from the exponent noted for steel erosion [14]. Therefore, the experimental results may be related to the process occurring in the real boiler of the Opole = 5.2 m/s is the same as that in the test stand (blowing nozpower plant under the assumption that the gas phase velocity w zle switched off). For the typical coal used in the power station, the mean concentration of fly ash in the flue gas is in the range of 15 g/m3. Near the banks of the steam superheaters, RKS 81E blowers manufactured by the Clyde Bergemann company are used in the BP 1150 boiler [20–22]. The rate of the lance travel, extending and retracting (1 blowing cycle), is sc bl = 628 s. The blower takes in approximately 800 kg of steam per cycle. The blower lance makes a rotary motion, during which steam jets from 3 nozzles blow away the ash deposits on the tube banks. The inside diameter of the blower lance is 80.9 mm. Each blower has 3 nozzles with a diameter of approximately 21.1 mm (Fig. 12), which gives a total steam outflow area of FL = 0.00105 m2. The steam pressure is approximately in the range of (0.6–1.0) MPa [21]. This type of pressure range involves a very high velocity of steam at the outlet of the nozzles. The steam jet blower flow characteristics: 800 _ bl ¼ smst ¼ 628 steam mass flow m ¼ 1:274 (kg/s) c bl
lance section area F l ¼
pd2l 4
2
¼ p0:0809 ¼ 0:00515 ðm2 Þ 4
superheated steam parameters at the blower inlet
p ¼ 1 MPa i ¼ 3264 ðkJ=kgÞ; q ¼ 3:26 ðkg=m3 Þ; s ¼ 7:465 ðkJ=kg KÞ t ¼ 400 C _
bl steam flow velocity in the blower lance w ¼ m ¼ 75 (m/s) Fq l
stagnation parameters of superheated steam for isentropic change s = idem
i0 ¼ i þ
w2 ¼ 3267 ðkJ=kgÞ ) 2
t 0 ¼ 401:5 C;p0 ¼ 1:01 MPa
.The equation of continuity implies that the maximum value, (wq) = max, determined by means of a trial method and subject to the condition of constant entropy s0 = sL, occurs at the smallest cross-section of the nozzle. The calculation results are listed in Table 8. The steam velocity at the nozzle outlet is calculated according to the following dependence [23]: Table 8 Characteristic parameters of the steam at the exit of the blowing nozzle. pL bar
iL kJ/kg
kg/m3
qL
wL m/s
wLqL kg/m2 s
9 8 7 6 5.4 5
3232 3197 3159 3116 3088 3067
3.012 2.740 2.475 2.198 2.024 1.908
265 373 465 549 598 632
798 1022 1151 1207 1212 1206
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489 Table 9 and bc ; bw . . Tube thickness loss versus mean flue gas velocity w Sootblower on bc = 1.2 bw = var Air p=3 Bar bl ðm=sÞ w 5.2 ðm=sÞ w 5.2
wL ¼
wmax ðm=sÞ Dh (mm)
Dh (mm)
Superheated steam p = 6.3
p = 10.5
24 40 70 0.1887 0.874 4.71 Sootblower off bc = 1.2; bw = 1.25 ðwmax ¼ 1:25 wÞ 0.00374
p = 20
p = 6 bar
84 8.14
100 13.6
qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2ði0 iL Þ
ð14Þ
The superheated steam leaves the nozzle with a velocity of wL 600 m/s at a pressure of pL = 5.4 bar when p0 = 10 bar (Table 8) pL = 3.3 bar when p0 = 6 bar. The difference between the pressure forces in the smallest section and the ambient pressure forces results in an increase of the stream momentum and, consequently, in an increase of the jet velocity outside the nozzle to the value w2.
w2 ¼ wL þ
FL ðp pam Þ105 _ bl L m
ð15Þ
If the pressure at the nozzle inlet is approximately 10 bar, the velocity resulting from expansion can reach values of up to 960 m/s (and w2 780 m/s at 6 bar). However, the jet of steam or air is slowed down rather abruptly at a short distance. The process is presented in Fig. 10 in relation to the air (perfect gas) flow for different values of the pressure [14]. Even after being slowed down by the start of the nozzles (the blower operation), the jet velocities are definitely higher than the velocity of the flow of dust-laden flue gases through the bank with shut down nozzles. The charts show that the higher the pressure before the nozzle, the higher the pressure after it. Thus, due to secondary expansion, the velocity of the gas agent in the area before the tube bank is also higher. The outermost tubes of the neighbouring stages of the steam superheaters, between which the steam jet ash blower is installed, are approximately 2 330 mm. Fig. 11 shows that the air jet velocity for y = 650 mm at a pressure of 10 bar is approximately 60 m/s, with a maximum velocity after secondary expansion of w2 = 510 m/s. In the case of the steam jet blower, where the velocity is w2 = 960 m/s after secondary expansion, the steam velocities before the tubes reach yet higher values. Therefore, for the analysis of the erosion caused by the temporary operation of the steam jet blower, the maximum values of the blowing agent are assumed to be at the level of wmax = 100 m/s. The erosion rate ve bl related to the use of blowers is unusually high, mainly due to the huge velocities given to the ash particles by the blowing agent. The factor that reduces the material loss is the relatively short annual time sbl a of the duration of the process described above compared to the annual operating time of the boiler. Below, a calculation of time sbl a is performed for the conditions prevailing in the area of the steam superheater of the BP1150 boiler. Due to the different dimensions of the test stand and the real boiler duct, the ratio of sootblowing stream in the total gas phase stream, which, in real conditions, consists of flue gas and (while sootblowing) steam, is much smaller for the actual at the level y = 650, as shown in Table 7, occurred boiler. Therefore, noticeable changes in the total gas phase velocity w significantly. Thus, in the test stand, while in the boiler, switching the nozzle on and off does not change the mean velocity w is lower for the test stand in contrast to the corresponding the parameter describing the velocity distribution bw ¼ wmax =w parameter calculated for the actual boiler and the operating sootblower. Nevertheless, the maximum velocity for both cases will only depend upon the operating parameters of the sootblower (air nozzle). According to [10], an abnormal boiler operation, i.e., without sootblowing, was assumed with factors of bw = 1.25 and bc = 1.20–1.37. From the start of the blowing cycle, bw changes dramatically because of the extremely high velocities of the cleaning medium, as shown in Fig. 11. Table 9 shows the tube thickness loss Dh (lm) according to Eq. (1) for a mean = 5.2 m/s, bc = 1.20 and bw = 1.25 at various blowing velocities bw. duct velocity of w Fig. 13 depicts the range of thickness loss for two extreme sets of flue gas parameters when the sootblower is not oper = 5.2 m/s, bw = var, bc = 1.20 and w = 10 m/s, bw = var i bc = 1.37. ating: w As shown in Table 9, the sootblower operation increases the thickness loss (Dh) by approximately 50 times at wmax = 24 m/s and by approximately 3640 times at wmax = 100 m/s compared with normal boiler operation without sootblowing. This analysis shows that the pressure of the blowing steam, which influences its velocity, has a strong effect on the erosion rate, in particular when the distance between the sootblower nozzle and the tube bank is small. Thus, the optimisation of the steam pressure is crucial for the proper operation of sootblowing systems. If the blower is started three times per 24 h, which is considered the highest frequency used in practice, within one year (which, for the Polish power plants, corresponds to approximately sa ffi 6000 h), the annual time of exposure is sbl a = 8.5 h [14].
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W. Wojnar / Engineering Failure Analysis 33 (2013) 473–489
¼ 10 m=s, bw = var and bc = 1.37 - - - - w ¼ 5:2 m=s, bw = var and bc = 1.20. Fig. 13. Tube thickness loss versus maximum velocity wmax —— w Table 10 and bc, bw. Tube thickness loss versus mean flue gas velocity w Sootblower on – sbl a = 8.5 h/a bc = 1.2 bw = var Air p = 3 bar bl ðm=sÞ w 5.2 ðm=sÞ w 5.2
bl ðm=sÞ w 10 wðm=sÞ 10
wmax ðm=sÞ Dh (mm)
Dh (mm) P Dh (mm)
wmax ðm=sÞ Dh (mm)
Dh (mm) P Dh (mm)
Superheated steam p = 6.3 bar
p = 10.5 bar
24 40 70 0.0016 0.0074 0.0396 Sootblower off – sn bl a = 5991.5 h/a bc = 1.2; bw = 1.25 0.0224 0.024 0.03 0.062 Sootblower on – sbl a = 8.5 h/a bc = 1.37 bw = var Air p = 3 bar p = 6.3 bar p = 10.5 bar 24 40 70 0.0021 0.01 0.053 Sootblower off – sn bl a = 5991.5 h/a bc = 1.37; bw = 1.25 0.213 0.215 0.223 0.266
p = 20 bar
p = 6 bar
84 0.069
100 0.116
0.091
0.138
p = 20 bar 84 0.092
Superheated steam p = 6 bar 100 0.155
0.305
0.368
It may be assumed that the boiler banks are exposed to a more intense erosion for 8.5 h of the 6000 h of operating time per year. With these assumptions, the tube material loss according to formula (1) is the total of the loss without blower operation (sn bl a = 5991.5 h) and the loss with blower operation for 8.5 h according to formula (3). Example calculation results are listed in Table 10. The steam expansion behind the nozzle is related to the lowering of the temperature. However, there is no danger of creating wet steam with water droplets, which might erode the heating surfaces of the boiler. In the worst case, assuming isentropic expansion from pL = 5.4 bar to ambient pressure pam = 1 bar, we obtain a final temperature of approximately 120 °C. For pL = 3.3, this temperature amounts to 170 °C. In actual conditions (irreversible process), the final temperature will be even higher. Furthermore, the blowing steam is introduced to the hot flue gas and sustains the steam in a superheated state. 5. Conclusions 1. The laboratory experiments and numerical modelling showed clearly that the mechanism of sootblower erosion consists of entraining the fly ash particles from the flue gas into the high-velocity blowing medium stream. Our calculations indicate that there is no danger of creating wet steam containing water droplets while steam expands and the tube wear as a result of the impact of water droplets flowing with the steam shall be neglected. 2. The conformity (with accuracy sufficient for most engineering purposes) between CFD modelling and experiments proves that the numerical modelling may be used as a reliable tool for a quantitative estimation of sootblower erosion in power boilers. 3. Approximate calculations showed that the sootblower erosion rate is many times higher than the fly ash erosion during a normal boiler operation. Therefore, even relatively short sootblower operation periods may lead to serious damages to the heating surfaces. 4. An important parameter of influence is the pressure of the sootblowing medium. Thus, the lowest pressure ensuring satisfactory cleaning effect and minimum thickness loss should be determined experimentally. References [1] Meuronen V. Ash particle erosion on steam boiler convective section. Lappeenranta University of Technology, Research Papers 64, Lappeenranta; 1997. _ [2] S´wirski J. Studies on ash erosion and the assessment tube erosive wear in boilers (Badania erozji popiołowej i ocena zuzycia rur kotłowych wskutek jej działania). Scientific books of the Institute of Power Engineering, Book 1, Warsaw; 1975 [in Polish].
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