Turbo Machinery Dynamics Part 10 pptx

(Gupta, Lewis, and Daurer, 2000). A 30° swirler for annulus # 1 and +50° and –50° swirlers for annulus # 2 are used to investigate flames produced from the change of swirl direction in the annuli. The arrangement with both swirlers having positive angles is referred to as producing coswirl flame, while angles of opposite directions have a counterswirl flame. Figure 9.22 provides a diagram of the burner outlet where the flame stabilization zone occurs, and Fig. 9.23 shows details of the swirling flow field and the regions of a swirl- stabilized premixed flame (Marshall, 1996). High-frequency temperature measurements are taken with a microthermocouple probe, with a wire diameter small enough not to cause interference on the flame’s structure while providing rigidity for the probe. At every location in the flame, the signal is amplified and digitized for a sampling time of 30 s to allow averaging over low-frequency temperature measurements and to assure a good statistical representation of the thermal field. Large variations in the temperature are present at any location in the flame. The sampling fre- quency used is 10 kHz, which is high enough to resolve small thermal time scales in the flame. Direct flame photographs taken during the tests provide data about the overall fea- tures of the flame and its stability. Negative images of the photographs determine the size of the flame in proportion to the burner. Raw temperature data have to be compensated for radiation losses and thermal inertia effects of the thermocouple. Radiation losses can be significant, particularly at high tem- peratures. Similarly, the level of fluctuations obtained without compensating the thermo- couple output can be considerable. Fluctuating temperatures are lower by as much as 250°C at some locations in the flame without compensation. Qi, Gupta, and Lewis (1997) provide a method for making corrections. The compensated mean temperature maps, shown in Fig. 9.24, display substantial differences between the left and right sides of the counter- swirling flame, with a flat hot shear layer present at the left side where temperatures exceed 1700 K. The shear layer on the right is steeper but comparatively cooler at about 1500 K. The nonsymmetric behavior of the counterswirling flame is observed by compar- ing it with the coswirling map. In the postflame region large differences exist in the mean temperatures on both sides. The coswirling map tends to be wider with a long area of COMBUSTION SYSTEM 343 FIGURE 9.22 Double concentric burner out- let region (Gupta, Lewis, and Daurer, 2000). reduced temperature fluctuations. A thin but intense reaction zone in the counterswirl case causes nonsymmetrical fluctuations in a smaller area of the flame. But overall differences in mean temperatures between the two cases are not large. The temperature maps make it possible to locate the combustion area, the recirculation zone and the postflame region. Outside the shear layers the flame tends to show higher 344 COMPONENT DESIGN FIGURE 9.23 Swirling flow field of premixed flame (Marshall, 1996). Product recirculation Postflame region rear stagnation point Recirculation zone boundary Recirculating fluid Shear layer Shear layer Environment Fresh reactant ignition Forward stagnation point Mixing reaction (with shear layer) Reactant burnout Recirculating zone A2 A1 A1 A2CP FIGURE 9.24 Compensated mean temperature maps (Gupta, Lewis, and Daurer, 2000). 0.0 2.0 1.5 1.0 0.5 0.0 2.0 1.5 1.0 0.5 Mean temperature counter (case 3a) Mean temperature counter (case 3cp) −1.0 −0.5 0.0 0.5 1.0 −1.0 −0.5 0.0 0.5 1.0 Coswirl Counterswirl Axial location (z/D) Axial location (z/D) Radial location (r/D) Radial location (r/D) 700 700 300 500 1300 1100 900 1500 1300 300 300 500 700 1300 1100 900 700 fluctuating temperatures than in the recirculation zone. The regions of low fluctuations are caused by continuous combustion, and represent pockets of burned gases within the recircula- tion zone. High temperature fluctuations outside the combustion zone are caused from mixing between the hot reaction products and the surrounding air, and suggest a stream of ambient air entrained toward the flame caused by the recirculation zone of the swirling flow field. Thus, the regions of high temperature fluctuations are outside the hot regions. Examination of the effects of swirl on the flame shape, mean and fluctuating temperatures can be useful in evalu- ating eddies present in the flame, which subsequently affect formation of NO x . 9.8 DRY LOW NO X COMBUSTION SYSTEM Cogeneration systems using a gas turbine as the prime mover offer high total thermal effi- ciency. They are subject to strict NO x regulations since air pollution in major population centers shows no sign of improvement. Water or steam injection or SCR is widely used in gas turbines to reduce the emissions, but the methods tend to increase the operating cost. Lean premixed combustion offers a convenient method to reduce NO x emissions with low initial and running cost. Tokyo Gas has focused on the development of dry low NO x combustors for a cogener- ation system in the 1 to 4 MW output range (Sato, Mori, and Nakamura, 1996). Engine out- put is controlled by varying the fuel gas flow, thus eliminating the need for complex variable geometries for air flow control. The double swirler staged combustor uses tertiary premix nozzles located around the liner. Multistaged combustion offers the benefit of sus- taining stable combustion with flame temperature in a range under 1650 K. Figure 9.25 shows the flow of air and gas in the double swirler combustor concept. COMBUSTION SYSTEM 345 FIGURE 9.25 Double-swirler-staged combustor arrangement (Sato, Mori, and Nakamura, 1996). Primary swirler Secondary swirler Secondary fuel Pilot fuel Primary fuel Tertiary fuel Primary nozzle Secondary nozzle Secondary port Air Exhaust gas Tertiary nozzles Pilot nozzlePilot swirler The primary and secondary premixing nozzles are placed coaxially with the radial swirlers. A pilot nozzle installed at the center of the premixing nozzles generates a diffu- sion flame rather than a perfect premixed flame, hence it stabilizes the flames at the other nozzles adequately. The swirlers generate swirling flows in the same direction. Four sepa- rate fuel lines lead to the pilot, primary, secondary, and tertiary nozzles. Figure 9.26 pro- vides the fuel supply schedule to the nozzles. The schedule is designed to provide constant air excess ratios for the pilot and primary nozzles over the whole engine operating load regime to generate stable combustion with low NO x emission. In the 0 to 30 percent load mode the schedule eliminates fuel supply to the tertiary nozzle to generate a lean fuel-air mixture with an excess air ratio of about 2.0 in the secondary nozzle, thereby igniting and oxidizing while directly contacting the stable combustion products of the primary and pilot nozzles. The flame temperature is low because of the excess air, producing practically no NO x . In the high engine load mode up to 100 percent fuel is supplied to the tertiary nozzle, fuel flow to the other nozzles remaining constant at maximum levels. Excess air ratio in the tertiary nozzle is also high to reduce NO x formation substantially. Operating conditions and target performance of the combustor are shown in Table 9.6. Target NO x level is 9 ppm at engine load between 50 and 100 percent, the normal operating range of gas turbines for cogeneration. At less than 50 percent load the target is 25 ppm. These target levels convert to 3.0 and 8.3 ppm under atmospheric pressure, assuming the general relationship that NO x emission is proportional to the square root of operating pressure. 346 COMPONENT DESIGN FIGURE 9.26 Fuel gas supply schedule (Sato, Mori, and Nakamura, 1996). Low mode 100 40 0 255075100 High mode Total fuel Primary fuel Engine load (%) Fuel ratio (%) Tertiary fuel Secondary fuel Pilot fuel TABLE 9.6 Operating Conditions and Target Performance Inlet air pressure 0.91 MPa Inlet air temperature 640 K Full load outlet gas temperature 1473 K Full load excess air ratio 2.7 Fuel, LHV Natural gas, 41.6 MJ/N⋅m 3 NO x target: 0–50% load <25 ppm (15% O 2 ) 50–100% load <9 ppm (15% O 2 ) Combustion efficiency target: 0–50% load >95% 50–100% load >99% Pressure drop 3% A schematic representation of the test facility is shown in Fig. 9.27. Air is preheated and rectified before introduction into the test combustor. The natural gas used in the test is 89 percent methane, with ethane, propane, and other hydrocarbons forming the rest. Combustion exhaust gas is sampled with a five-point water-cooled probe positioned down- stream of the exhaust, then introduced into the gas analyzer. O 2 is analyzed by a magnetic analyzer, CO and CO 2 with a nondispersive infrared analyzer, NO x with a chemilumines- cence analyzer and UHCs with a flame ionization detector. Temperature distribution at the combustor outlet is measured in the same plane as the gas-sampling probe at 24 positions to obtain an acceptable pattern factor. Total flow rate of the process air is calculated from exhaust gas composition and measured fuel flow, and flow rate to each nozzle assumes a split proportional to the open area of the respective air nozzle. The charts of Fig. 9.28 provide performance characteristics of the combustion system, showing the effects of excess air ratio and corresponding engine load on NO x , CO, UHCs, and combustion efficiency. Engine loads of 100, 50, and 0 percent are associated with 2.7, 3.9, and 6.8 of excess air ratio l tot . The NO x level is considerably influenced by the high/low engine load. In the low mode, when l tot is between 5.0 and 7.0, the NO x holds steady at 5 ppm. The secondary flame produces virtually no NO x in this range. But a sharp increase is observed when l tot decreases from 5.0 to 4.0, with NO x level reaching 8 ppm. In the high mode, the additional tertiary fuel with its sufficiently high excess air produces lesser NO x , dipping under 2 ppm when l tot is 3.3. Thus, a higher excess air ratio helps to curtail thermal NO x production. Combustion efficiency reaches a low value of 95 percent during the low mode engine operation for l tot = 6.2, with a corresponding increase in UHC formation. The efficiency curve recovers, as excess air ratio reaches a maximum. Combustion efficiency and CO and UHC emissions at maximum l tot are affected by the combination of pilot and primary COMBUSTION SYSTEM 347 FIGURE 9.27 Schematic diagram of test facility (Sato, Mori, and Nakamura, 1996). TV camera Water Exhaust Silencer Spray Cooling tower Heat exchanger Blower Preheat burner Test combustor Natural gas Gas flow meters Combustion gas sampling probe Thermocouples M M M M 348 COMPONENT DESIGN FIGURE 9.28 Combustor performance characteristics (Sato, Mori, and Nakamura, 1996). FIGURE 9.29 Double-swirler-staged combustor design (Sato, Mori, and Nakamura, 1996). f 210 f 290 633 flames. A similar drop in combustion efficiency occurs in the high operating engine mode when l tot increases from 4.0 to 4.6, where the tertiary flame plays the same role as the sec- ondary flame does in the low mode. Considering that the pilot burner is designed for a sta- ble diffusion flame, the majority of CO and UHC is originated in the primary flame. Unlike the pilot, the primary flame mixes directly with the secondary airflow, so the swirling pri- mary flame has a limited residence time and results in high emissions of CO and UHC. The relatively simplified geometry of the double-swirler-staged combustor designed to operate at standard atmospheric pressure is shown in Fig. 9.29. 9.9 CATALYTIC COMBUSTOR FOR UTILITY TURBINE The SCR method is useful in chemically reducing NO x to nitrogen and water vapor; how- ever, the costs associated with heat rate deterioration due to diluent injection and the capi- tal and operating costs for the required systems make it financially unattractive for application in combined cycle and cogeneration power plants incorporating gas turbines. Direct catalytic combustion offers good potential for reducing formation of NO x , CO, and UHC in tests carried out at General Electric for model MS9001E gas turbine (Dalla Betta et al., 1996; Schlatter et al., 1997). The design calls for partial reaction of fuel-air mixture within the catalytic reactor to generate a gas temperature of 800–1000°C at reactor exit. At this temperature in the reactor, the catalyst can include precious metals, and the substrate may be cordierite or metal. The combustion system design (Fig. 9.30) requires a preburner, fuel and air preparation system, catalytic reactor, and a combustion liner downstream of the reactor. The preburner carries machine load at conditions when temperature levels do not allow satisfactory cat- alytic combustion, and also preheats to achieve catalytic reactor ignition at high loads. Catalytic staging initiates at turbine inlet temperature of 700°C when the main fuel injec- tor activates. The fuel and air preparation system provides the components and preburner products to the reactor bed at a uniform strength, pressure, velocity, and temperature. COMBUSTION SYSTEM 349 FIGURE 9.30 Catalytic combustion test rig (Schlatter et al., 1997). Preburner fuel inlet Preburner Main fuel inlet Main fuel injector Catalyst Video camera Postcatalyst reaction volume Transition piece Nozzle box (turbine inlet) Air inlet Perforated plate The catalytic reactor promotes oxidation of hydrocarbons and CO for lean mixtures at adiabatic flame temperature below the threshold for thermal NO x formation. Combustion initiated by the catalyst is then completed by homogeneous burning in the postcatalyst region where high temperatures are obtained. Catalytic reactor technology developed by the manufacturers gives a bed for full fuel and airflow required for maximum power, while avoiding exposure of the catalyst to high temperatures that may damage the supporting sub- strate. Use of ceramic catalytic materials maintains the catalyst surface below the adiabatic combustion temperature. Advantage is taken of the palladium oxide in catalyzing methane oxidation, while metallic palladium is appreciably less active (McCarty, 1994). Palladium has the unique thermodynamic characteristic of oxidizing and reducing. Depending on pressure, the oxide decomposes to the metal between 780 and 920°C. The reactor consists of three separate catalyst stages, with the stages formed by corrugating and foiling metal foil to constitute a channeled monolithic structure. Active ceramic material is coated on the foil. The stages are supported in a reactor container by large cell honeycomb structures made of Hastelloy X. Experimental data are obtained over a range of conditions from full speed without load to base load of the engine. Combustor discharge temperature ranges from 543°C at no load to 1195°C at base load. Reactor operation is started by heating the system with the pre- burner, then turning on the main fuel flow to provide a smooth light-off of the reactor with a uniform temperature profile across the face. Figure 9.31 shows measured pollutant emissions data corresponding to ISO ambient with 15 percent oxygen concentration for average combustor temperatures at the nozzle box. The peak NO x value of 55 ppm in the 519–626°C range results from the diffusion flame in the preburner when no fuel is delivered to the main burner and the catalyst. As the combustor is taken to higher exit temperatures, fuel is shifted to the main fuel injector. NO x levels drop to a less-than-desirable 11 ppm, mostly due to the need for higher temperature in the preburner to keep the catalytic reactor fully active. Introduction of steam into the pre- burner zone lowers the NO x to 3 ppm. These data are consistent with existing data for NO x suppression by steam injection for diffusion combustors using natural gas (Touchton, 1984). At base load the catalytic reactor fuel is about 80 percent of the total, indicating essentially no NO x production by the reactor. CO emission shows a similar trend, peaking to 3200 ppm at 930°C during preburner only operation, when catalyst staging is in a transient condition between no load and base load. 350 COMPONENT DESIGN FIGURE 9.31 NO x , CO, UHC Emissions (Dalla Betta, 1996). Combustion temperature rise subsequently transfers to the catalytic combustor. At base load the reaction temperature is noted at 1196°C, when CO emission falls to a minimum value of 10 ppm. Preburner exit temperature must be maintained at a high enough level to keep the catalyst fully active. Considerable scatter is noted at the base point, mostly due to sensitivity to the preburner exit temperature. UHC emissions show a major peak at 800°C during preburner fuel operation, reducing to negligible levels at 1200°C. At the simulated base load operating point with preburner exit temperature at 563°C, the overall combustion of fuel to equilibrium combustion products is greater than 99.99 percent. Dynamic pressure measurements indicate that the catalytic combustor system experi- ences oscillations lower than in conventional combustors. The maximum discrete peak has a magnitude of 0.00173 MPa at a frequency of 252 Hz, and occurs during steam injection. Maximum overall root-mean-square (rms) noise level of 0.00836 MPa also occurs at the same time. Without steam injection the dynamic pressure measures about 20 percent less. Combustor exit temperature distribution factor, defined by the ratio of maximum variation from the mean to the overall combustor temperature, is 0.138. Preburner exit temperature nonuniformity contributes substantially to this variation. The tests indicate improvement in the structural integrity of reactor, with the diameter experiencing minimal distortion after several hours of operation. 9.10 ACOUSTIC RESONANCE Combustion of the air and fuel mixture is accompanied by noise directly as a conse- quence of the process and indirectly due to the flow of burned gases through the turbine and exhaust nozzle. Combustion noise can become detrimental when instabilities aris- ing in the burning process couple with acoustic modes inside the chamber. The natural frequencies of the combustor can be excited by resonant pressure waves in the main gas flow along the axial and radial directions, as also by lateral modes in the tangential direction (Paxson et al., 1995; Ohtsuka et al., 1998). Sustained oscillating phenomena due to a higher level of mixing of the fuel and air prior to combustion lead to engine noise and vibration problems. Premixed combustion in gas turbines helps produce low levels of NO x emissions, but practical application of this concept is limited by self-excited combustion oscillations. When operation in a lean, premix combustor is close to the flammability limit, slight changes in operating conditions can lead to sudden flame extinction or to excessive CO emissions. In addition to static stability, lean premix combustors must achieve dynamic sta- bility, meaning the combustion must not oscillate. Oscillation must be eliminated in a com- bustor design because the associated pressure oscillations tend to have life shortening consequences (Richards and Janus, 1997). Figure 9.32 shows cracks experienced in a tran- sition piece due to excessive acoustic oscillations. Operation near the lean limit is especially prone to oscillation problems, where minor variations in fuel-air ratio lead to appreciable variations in combustion reaction rate. When these variations in the reaction rate couple with the acoustic modes, significant pressure oscillations occur, with frequencies ranging from hundreds to a few thousand Hertz. The task of studying and eliminating combustion oscillations in a gas turbine is com- plicated by the specific acoustic response of a combustor’s design. The combustion process interacts with the acoustic field, leading to instabilities. Rapid changes in air and fuel sup- ply and aerodynamic disturbances may lead to the instability because of a sequence of extinction and reignition of the flame in parts of the combustor. If the heat release rate does not take place uniformly and periodic spikes occur, acoustic waves of the same frequency may be expected in the combustion zone. Reflection from the liner causes pressure waves COMBUSTION SYSTEM 351 to be returned to the combustion zone after a time delay, and the waves are reinforced when the heat release and pressure wave peaks coincide. As defined by Lord Rayleigh’s criterion, oscillations set in when changes in heat release are in phase with acoustic pressure distur- bances. Conversely, oscillations are dampened when heat-release fluctuations are out of phase with pressure fluctuations. This criterion serves as the cornerstone for the develop- ment of combustion oscillation analysis. Variation in heat release results from changes in flame structure produced by acoustic pressure disturbances. Time delay between pressure disturbance and heat-release variation determines the phase and, consequently, the stabil- ity of the system. Based on these observations, lean premix combustors can be characterized by a simple time-lag approach. Figure 9.33 shows for a specific case a schematic diagram of the impor- tant processes, where a sinusoidal pressure disturbance produces a sinusoidal variation in airflow 180° out of phase with the pressure. 352 COMPONENT DESIGN FIGURE 9.32 Acoustic oscillations damaged transition piece (Lieuwen and McManus, 2002). FIGURE 9.33 Flow characteristics during acoustic oscillation (Richards and Janus, 1997). [...]... temperature near the wall, and is expressed by C1 = 0 .10 × x  ka  (Re x )0.8    s  x −0.36 (Tw,ad − Tw1 ) (9 .10) Hence 0.153  0.036  C1 = 0 .10 ×  × (297,640)0.8 ×   0.036   0.002  −0.36 (1209 − Tw1 ) = −1287 × Tw1 + 1, 556, 038 From Probs 9.1 through 9.4, 2 R1 = 765,126 − 0.003066 × Tw15 W/m 2 4 R2 = 2.0934 × 10 −8 × Tw 2 − 14, 997 W/m 2 C2 = 100 3 × Tw 2 − 922, 719 W/m 2 K1−2 = 25, 000 × (Tw1... R  P  c 2 (10. 9) where static pressure P = W/LD, and L is the bearing length The friction coefficient, the ratio between frictional (Fτ) and bearing loads (W), is f = Fτ W (10. 10) The general shape of the coefficient of friction in terms of the Sommerfeld number is given in Fig 10. 4 For low values of S the friction factor tends to be high because of boundary condition effects, when partial contact... for the thermal aspects of the problem q dq   ∆T = −a (T − T1 ) = ln 1 + E ∫  q1 G h2 t   q dq m1 = 1 + E∫ q1 G h 2 m t (10. 5) (10. 6) where E is an adiabatic parameter Typical values for the turbulence coefficients are provided in Fig 10. 2 BEARINGS AND SEALS FIGURE 10. 2 375 Turbulence coefficient values Hydrodynamic pressures start developing at the downstream edge of the oil feed groove, and... Turbines & Power 104 :377–385, 1982 Yoshida, Y., Oyakawa, K., Aizawa, Y., and Kaya, H., “A high temperature catalytic combustor with starting burner,” ASME Paper # 00-GT-087, New York, 2000 BIBLIOGRAPHY Darling, D., Radhakrishnan, K., Oyediran, A., and Cowan, E., Combustion-Acoustic Stability Analysis for Premixed Gas Turbine Combustors, NASA TM 107 024, 1995 CHAPTER 10 BEARINGS AND SEALS 10. 1 INTRODUCTION... for journal bearings is given by E= FIGURE 10. 3 2µ1αω  R  bw  c  2 Temperature and pressure profiles in bearing lubricant film (10. 7) 376 COMPONENT DESIGN E is thus a function of lubricant properties and conditions of bearing operation For a wellaligned journal, the film thickness for eccentricity ratio ε = e/c is defined by h/c = 1 + ε × Cos(θ − φ) (10. 8) Minimum film thickness is obtained when...       (10. 2) 374 COMPONENT DESIGN FIGURE 10. 1 Fluid film journal bearing where q and z are the circumferential and axial coordinates, Gq and Gz are the turbulence coefficients in the two directions and are functions of the Reynolds number, h is the film thickness and is a function of q and z, R is the journal radius, m is the lubricant viscosity, and p is the pressure In Fig 10. 1 the arc from... (Taverage − T1 ) = H bwQz (10. 12) When thermal effects are included the variable temperatures may be combined by the factor correlating temperature with viscosity The evaluation of the bearing performance parameters for specific operating conditions can then be obtained for the desired bearing geometry FIGURE 10. 4 Friction factor in fluid film bearings BEARINGS AND SEALS 377 10. 3 JOURNAL BEARING TYPES... TYPES Many different types of bearings are used to support a rotating shaft Some representative types are shown in Fig 10. 5 The partial arc and the grooved bearings are adaptations of the plain cylindrical bearings With a fixed geometry a cylindrical bearing tends to be an FIGURE 10. 5 Journal bearing types 378 COMPONENT DESIGN unstable form of support The elliptical, three-lobed, and offset cylindrical... role of Tmax Circular bearings have the advantage of ease in manufacturing, installing, and repair Many configurations are available for turbomachines, with the designs based on a number of arcs to form a circular geometry The differences arise from the number of partial arcs and in the relative location of the centers of curvature of the arcs and of the assembled bearing Circular bearings may be provided... 0.05 m Liner cross-sectional flow area Aa = [p × (0.22 – 0.152)]/4 = 0.01374 m2 Combustor inlet temperature T = T3 = 920 K Then 0.8 C2 = 0.020 × 0.0553  7.0  (T − 920) w2 0.050.2  0.01374 × 3.85 × 10 5  = 100 3 × Tw 2 − 922480 W/m 2 Conduction heat transfer through a solid liner wall due to a temperature gradient in the wall K1−2 is K1−2 = = kL (T − T ) t L w1 w 2 25 (Tw1 − Tw 2 ) 0.001 = 25000(Tw1 . location (z/D) Radial location (r/D) Radial location (r/D) 700 700 300 500 1300 1100 900 1500 1300 300 300 500 700 1300 1100 900 700 fluctuating temperatures than in the recirculation zone. The regions. COMPONENT DESIGN FIGURE 9.26 Fuel gas supply schedule (Sato, Mori, and Nakamura, 1996). Low mode 100 40 0 25507 5100 High mode Total fuel Primary fuel Engine load (%) Fuel ratio (%) Tertiary fuel Secondary. MJ/N⋅m 3 NO x target: 0–50% load <25 ppm (15% O 2 ) 50 100 % load <9 ppm (15% O 2 ) Combustion efficiency target: 0–50% load >95% 50 100 % load >99% Pressure drop 3% A schematic representation