热斑对涡轮级影响

热斑对涡轮级影响
热斑对涡轮级影响

NUMERICAL INVESTIGATION ON UNSTEADY EFFECTS OF HOT STREAK ON FLOW AND

HEAT TRANSFER IN A TURBINE STAGE

Bai-Tao AN

Institute of Engineering Thermophysics Chinese Academy of Sciences

Beijing, 100080,China

E-mail: anbt@https://www.360docs.net/doc/e94407715.html,

Jian-Jun LIU

Institute of Engineering Thermophysics Chinese Academy of Sciences

Beijing, 100080,China

E-mail: jjl@https://www.360docs.net/doc/e94407715.html,

Hong-De JIANG

Department of Thermal Engineering

Tsinghua University

Beijing, 100084,China

E-mail: jianghd@https://www.360docs.net/doc/e94407715.html,

ABSTRACT

This paper presents a detailed flow and heat transfer cha-racteristic analysis on gas turbine first stage turbine under hot streak inlet conditions. Two kinds of inlet total temperature conditions are specified at the turbine stage inlet. The first is uniform inlet total temperature, and the second is hot streak 2D total temperature contour. The two kinds of inlet conditions have the same mass-averaged total temperature, and the same uniform inlet total pressure. The hot streak total temperature contours are obtained according to the exit shape of an annular-can combustor. The ratio of the highest total temperature in the hot streak to the mass-averaged total temperature is about 1.23, and one hot streak corresponds to two vane passages and four blade passages. Six hot streak circumferential positions relative to the vane 1 leading edge varied from -2% to 81% pitch are computed and analyzed. Results show that hot streak obviously increases the non-uniform degree of vane heat load in compari-son with the uniform total temperature inlet condition. The change of hot streak circumferential position leads to the cir-cumferential parameter variation at stator exit, and also leads to different transient periodic fluctuating characteristics of heat load and pressure on rotor blade surface. The hot streak of rela-tive pitch at 65% obtains similar heat load for the two vanes corresponding to one hot streak and small fluctuation of the averaged heat load on rotor blade.

Keywords: unsteady flow, hot streak, three dimensional numer-ical simulation, gas turbine INTRODUCTION

The inlet temperature field of gas turbine is non-uniform or distorted usually, as the combustor exit flow has both the cir-cumferential and radial temperature gradients. These tempera-ture gradients arise from combustor framework characteristic and combustor surface cooling. Hot streaks have temperature typically twice the free stream temperature in aero engine. For the heavy duty gas turbine, the ratio of the highest temperature to mass-averaged temperature is slightly low. It is well known that the most important problem in gas turbine is blade or vane reliability under excessive working temperature. The reliability includes two aspects, steady heat reliability such as cooling design of stator vane, and unsteady heat reliability such as thermal fatigue of rotor blade. Hot streak influences the above two aspects greatly. Therefore, for gas turbine cooling design, especially the first-stage, the effect of hot streak must be consi-dered.

A number of studies have been performed for the hot streak effects on turbine stage. Jenkins et al [1] studied the film cooling effects on the dispersion of a simulated hot streak on experimental cascade. The result showed that film cooling re-duces the peak temperature values obviously. Jenkins and Bo-gard [2] performed experimental studies to discuss the attenua-tion of a simulated hot streak in the experimental vane cascade, and the major topic focused on the effects of turbulence level and pitch position variations of the hot streak to the peak hot streak temperature at a few axial positions. The studies indi-cated that the attenuation rate of hot streak depends on the loca-

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tion of the hot streak relative to the vane due to isolating of the hot streak core by the vane wall. Zhao et al [3] employed the numerical method to investigate the influence of the hot streak temperature ratio in a high pressure stage of a vaneless counter-rotating turbine. The temperature ratio of a circular shape hot streak varies from 1.0 to 2.4. The results showed that the heat load of the high pressure turbine (HPT) rotor increases with the increasing of the hot streak temperature ratio. Ji et al [4] also performed the studies on a counter-rotating turbine use 3D un-steady Euler solver, and the results showed that the effect of the hot streak on a counter-rotating turbine is nearly the same as a conventional turbine, however, clocking between the hot streak and the vane of the HPT exists significant influences on the heat load of the whole HPT stage. Dorney and Gundy-Burlet [5] studied a 1-1/2 stage turbine geometry, in which the posi-tion of the hot streak is varied to produce different levels of impingement upon the first-stage stator. The results indicated that when the hot streak impinges on the first stage stator, the pressure surfaces of the downstream blade rows reach much higher time-averaged temperature. He et al [6] carried out com-putational studies on the effects of hot streak circumferential length-scale in transonic turbine stage, both aero-force and aero-thermal are considered, and the results showed that the circumferential wavelength of the temperature distortion can significantly change the unsteady forcing as well as the heat transfer to rotor blades. Dorney et al [7] performed a numerical investigation of multi-blade count ratio and 3D effects on the hot streak migration in a turbine stage. Three-stator/four-rotor and one-stator/one-rotor were investigated using 2D and 3D numerical method. Gundy-Burlet and Dorney [8] performed a 3D unsteady simulation to investigate the effects of radial loca-tion of the hot streak in a 1-1/2 stage turbine, in which the radi-al location and the clocking position of the hot streak is various. Shang and Epstein [9] explored the influence of inlet hot streak temperature distortion on turbine blade heat load in a transonic axial flow turbine stage using a 3D multi-blade row unsteady Euler code, and found that temperature distortion significantly increases both blade surface heat load non-uniformity and total blade heat load by as much as 10-30 percent (mainly on the pressure surface). Zilli et al [10] discussed the performance of a single stage axial flow HPT, and the analysis demonstrated that the temperature distortion does not have a significant effect on the performance of the HPT apart from a small reduction in efficiency.

From the studies have been performed, for the heat load, more attentions are put on the variation of time-average value, few studies pay attention to the transient heat loads variation. The emphasis of the present work is on the effects of hot streak on flow and heat transfer characteristics of the first-stage stator and rotor. Six circumferential positions of the hot streak are considered. For the stator, it aims at exploring the hot streak migration in vane passages, exit circumferential parameter var-iation, and the effects of the hot streak on vane heat load. For the rotor, it aims at exploring the periodic transient variation characteristics of heat load and aerodynamic force under six hot streak configurations.

NOMENCLATURE

C Axial blade chord [m]

H Blade height [m]

h Wall heat transfer coefficient [W/m2.K]

h0Reference wall heat transfer coefficient, [W/m2.K] k Turbulent kinetic energy [m2/s2]

LE Leading edge (minimal x position)

Ma Mach number

N Time step number

p Pressure[Pa]

PS Pressure surface

q Wall heat flux [W/ m2]

q0Reference wall heat flux [W/ m2]

RPLE Relative pitch to vane 1 LE

SS Suction surface

TE1Trailing edge of vane 1

TE2Trailing edge of vane 2

t Pitch [m]

T Temperature [K]

Tu Turbulence intensity [%]

x Axial distance [m]

y+Non-dimensional distance from wall

Greek Symbols

αExit flow angle [degree]

ωDissipation rate of turbulent kinetic energy [m2/s3] Subscripts

0Inlet plane

Superscripts

*Stagnation condition

—Averaged value or relative value

GEOMETRY OF THE TURBINE STAGE

The first stage turbine model of a heavy duty gas turbine is shown in Fig.1. The span/chord ratio of the vane and blade is 1.1 and 1.9, respectively. The diameter/height ratio of the vane and blade is 18.6 and 17.8, respectively.

(a) Meridian view

(b) S1 view (N=0)

Fig.1 Model of the first-stage turbine

The turbine stage model includes stator domain and rotor domain. The stator-rotor interface is set at 1/2 location between the stator and rotor domain. There are 20 combustors, 40 vanes, and 82 rotor blades in the full annulus respectively in the actual

engine. For the transient simulation, the model is simplified to one combustor, two vanes, and four blades to assure the pitch ratio of the computational domains is 1:1:1, i.e., one hot streak corresponds to two vane passages and four blade passages. One hot streak periodicity has 18 degree circumferential angle. The vane number is labeled as vane 1 and vane 2 for analysis. Rotor blade tip clearance is not considered in the present simulations. 40 time steps are set for one rotor domain period, each time step is 2.5×10-5 seconds.

NUMERICAL METHOD AND VALIDATION

The 3D fully implicit coupled N-S solver, CFX5, is em-ployed to obtain the unsteady time accurate solution. This solv-er is based on finite-volume method, and the discretization scheme is second-order accurate. In the present simulation, k-ω turbulence model and structured grid are employed to solve the 3D unsteady viscous flow. Sliding mesh is adopted at the do-main interface. Fig.2a shows the grid distribution of the compu-tational domain. There are 0.42M and 0.4M grid elements in stator domain and rotor domain respectively, total 0.82M grid elements in the turbine stage. Grid is dense near the wall for assuring the accuracy of wall viscous flow and adapting the requirement of turbulence model, as shown in Fig.2b. General-ly, for k-ω model, the minimum wall y+ should be below 1.0. CFX also introduces automatic wall treatment to k-ω model to ensure the solution robustness. In the present study, the minim-al y

+ of the first layer grid is about 0.6.

(a) Domain grid

(b) Near wall grid

Fig.2 Computational grid for the present simulations

Steady solution is obtained before transient simulation starts. The steady solution is set as the initial flow for unsteady transient simulation. When performing the transient simulations, one monitor point is set in the rotor domain, the monitored pa-

rameter is temperature. Multiple periods are computed until the

rotor domain

To validate the accuracy of CFX and the applicability of k-ω turbulence model, a steady 3D simulation is performed to compare with the experimental data offered by Hylton [11]. The cascade chosen for comparison is Mark-II. Many experi-mental conditions have been performed in Hylton’s experi-ments. The experimental condition chosen here is No.5411. Figure 4 shows the computational grid of Mark-II cascade, the size of the grid elements has two cases for the verification of grid dependence. Case 1 has 0.55M grid elements, and case 2 has 0.20M grid elements. The density of near wall grid of case 2 is just like the setting in Fig.2b. The case 1 has denser grid in the flow direction and the radial direction especially near the wall than case 2. In the simulations, inlet total pressure, inlet total temperature, and exit static pressure is specified according to the test conditions. In addition, the wall temperature is also specified according to the test result. The computed wall heat transfer coefficient is compared with the test result, as shown in Fig.5.

It can be seen that the wall heat transfer coefficient of CFX solution agrees well with the experimental result whi-chever trend and value, except in the region from the leading edge to x/C =0.4 on the suction surface, the value of CFX solu-

tion is obviously higher than the experimental data. This may be caused by the unaccurate prediction of turbulence transition. In this case, SST turbulence model could be better, but SST model needs even fine grid near wall, and the computational time will be extremely long for transient simulations. In gener-

al, CFX code with k-ω turbulence model can satisfy the need of the present study, at least for qualitative analysis. In addition, the result difference between the fine grid (0.55M) and the rela-tive coarse grid (0.20M) is negligible.

BOUNDARY CONDITIONS AND SCHEMES

In the present simulation, uniform total pressure, uniform or non-uniform total temperature and uniform flow angles are specified at the stage inlet plane, and averaged static pressure is specified at the stage outlet plane. The inlet turbulence intensity is set to Tu=10%, and the turbulence length scale is 115mm. The rotation speed of the rotor blade is 3000rpm. Constant temperature about 0.8 T0* is set to the vane, blade and endwall referring to the temperature limit of blade alloy. Wall heat flux q is computed for the analysis of wall heat transfer characteris-tic, reference wall heat flux q0 is employed to obtain non-dimensional wall heat flux q/q0, and q0 is the average of q over

tribution along radial direction at stage inlet plane

To study the influence of inlet hot streak on flow and heat transfer in turbine stage, two kinds of inlet total temperature conditions are specified at stage inlet. The first is of uniform distribution of inlet total temperature, used for comparison pur-pose mainly. The second is 2D contours of inlet total tempera-ture to simulate the exit temperature field of combustor, which is hot streak inlet condition. For the hot streak inlet condition, the ratio of the highest total temperature to the mass-averaged total temperature is about 1.23, and the radial position of the highest total temperature locates at about 65% inlet height ac-cording to the actual turbine inlet condition. Circumferential mass-averaged total temperature distribution along the radial direction is shown in Fig.6. The two kinds of total temperature inlet conditions have the same mass-averaged total temperature at inlet plane.

To study the effects of the circumferential position varia-tions of the hot streak, the hot streak inlet condition has six schemes, as shown in Fig.7. The circumferential position of the highest total temperature core point changes from the leading edge of vane 1 to suction surface of vane 2, two adjacent schemes of the hot streaks have 1.5 degree circumferential in-terval as shown in Fig.7a. The black lines in Fig.7b denote the leading edge line of vane 1, the leading edge line is minimal x position of the vane but not actual stagnation line. The relative pitch of the hot streaks to the leading edge line of vane 1 (RPLE) varies at -2%, 15%, 31%, 48%, 65%, 81% pitch posi-tions.

(a) The circumferential positions of the inlet hot streak and

probing planes and points in stator and rotor domain

(b) Inlet total temperature contours

Fig.7 Inlet total temperature conditions

RESULTS AND ANALYSIS

Hot streak development in stator

Development process of the hot streak in stator vane pas-sage is very important for flow and heat transfer not only to vane but also to rotor blade. Figure 8 shows the relative posi-tions in the blade-to-blade surface between the hot streaks and the vane 1 leading edge. Six pitch positions of the hot streak can be divided into four cases. In the first case the hot streak locates at the leading edge of vane 1 as RPLE=-2% and 15%. In the second case the hot streak is close to the vane 1 pressure surface as RPLE= 31%. In the third case the hot streak locates at the middle passage between vane 1 and vane 2 as RPLE= 48%. In the fourth case the hot streak locates close to the suc-tion surface of vane 2 as RPLE=65% and 81%. Because one hot streak period faces two stator vane passages, when the six hot streak positions turn from the passage composed by vane 1 and vane 2 to next passage, the situation of vane 1 and vane 2 interconvert.

RPLE=-2% RPLE=15% RPLE=31%

RPLE=48% RPLE=65% RPLE=81%

Fig.8 Temperature contours at 65% vane height

The hot streak migrations in the two stator vane passages for the six relative pitch positions are shown in Fig.9. The high-est temperature region of RPLE=-2% situates at the leading edge of vane 1. When entering the vane passages, the elliptoid high temperature region is divided into two parts by vane 1. Along with the development downstream, the elliptoid shape is resumed quickly at suction side, whereas, at the pressure side the high temperature region is pressed to the wall and expands along the radial direction. At vane exit, the two sides of the vane 1 wake have different temperature shape obviously. At the suction side, high temperature region becomes narrow in the radial direction and wide in the circumferential direction, which is caused by the opposite pressure gradient and extrusion of 3D separation line on the suction surface. At the pressure side, high temperature region becomes very narrow in the circumferential direction and extend greatly along the radial direction com-pared with that of inlet. The RPLE=15% has the similar beha-vior to RPLE=-2%, a little difference is that a larger part of the high temperature core region situated at the vane 1 pressure side. This case is similar to the experimental inlet condition in Jenkins et al [1], although the flow conditions are not exactly the same. The temperature reduction of the hot streak at the suction side is faster than the pressure side within the vane pas-sage, i.e., x/C=20% to 80%, this also agrees with the tempera-ture reduction trend in Jenkins et al [1]. In this case, at vane exit, the high temperature region by sides of the wake has al-most the same size and temperature value due to flow mixing, but the radial location of the peak temperature is different, i.e., the radial location at suction side is higher than at pressure side obviously, which is caused by 3D viscous flow and meridional endwall curves of the stator together. When the hot streak flow very close to the vane wall, it will be influenced by the wall limiting stream line, see Fig.10. At the suction side, 3D separa-tion line caused by the 3D viscous flow exists at the rearward section. For the fluid near the wall, 3D separation line plays the role similar to constrict the meridional shape. At the pressure side, the wall limiting stream line varies with the meridional endwall curves, whereas, the meridional hub curve is diffused, so the radial position difference between suction side and pres-

sure side appears.

RPLE=-2% RPLE=15%

RPLE=31% RPLE=48%

total temperature inlet condition

The highest temperature region of RPLE=31% is close to vane 1 pressure surface at stator inlet, so it always transports along the pressure surface. At vane exit, it becomes closer to the pressure surface and has the same pressure-side-behavior with RPLE=-2% and 15%. The RPLE=48% behaves just like RPLE=31%, although the highest temperature region situates at

the middle position between vane 1 and vane 2 at stator inlet, the highest temperature region situates close to vane 1 pressure surface at the vane exit.

The RPLE=65% is similar to the RPLE=81%, i.e., the highest temperature region close to vane 2 suction surface at stator inlet. The developments of the hot streak have two phas-es. In the forward section of the vane passage, high temperature region is elongated along the radial direction corresponding to accelerated flow. In the rearward section of the passage, high temperature region extends along the circumferential direction corresponding to decelerated flow. At vane exit, temperature

across the highest temperature point (N=10)

Fig.11 and Fig.12 are the circumferential and radial tem-perature distributions at the position across the highest tempera-ture core point at stator exit plane. The position change of the highest temperature region through the vane passage is very clear. First, Fig.11 shows that the exit circumferential interval of the highest temperature core point between the adjacent hot streak schemes becomes disaccord with that of inlet. The cir-cumferential interval is small (<1.5 degree) close to the TE1 pressure side, the circumferential interval becomes large (>1.5 degree) gradually with the hot streak turning from TE1 pressure side to TE2 suction side. It should be noted that the exit circum-ferential position of the highest temperature core point for the case of RPLE=65% situates at the middle of TE1 and TE2. Second, Fig.12 shows obviously that the highest temperature core point for RPLE=-2% has the highest radial position than the other hot streak schemes, RPLE=81% behaves just like RPLE=-2%. Although the hot streak RPLE=-2% is classified to the case of impinging upon the vane 1 leading edge, from Fig.11 we can see that the most part of the high temperature region situates at vane 1 suction side, so this case can also be classified as the hot streak very close to the vane 1 suction sur-

face (i.e., RPLE=81%). From this point of view, this phenome-non confirms the analysis associated with Fig.10. The conclu-sion is that the radial positions of the highest temperature re-gion of the hot streak at stator exit have the increasing trend from the TE1 pressure side to TE2 suction side. But the highest location in radial direction cannot exceed the 3D separation line at stator exit. Third, at vane exit, small peak temperature differ-ence exists among the hot streak schemes. Jenkins and Bogard [2] indicated that the peak hot streak temperature at vane exit was the same for an impinging and non-impinging hot streak under the steady and adiabatic experimental conditions. The present result has the similar trend. The small peak temperature difference comes from two aspects. One is the heat transfer from the flow to the wall (isothermal wall instead of adiabatic wall). Another is the unsteady rotor-stator interaction of the downstream rotor blades. The RPLE=31% has the highest peak temperature for N=10 mainly due to the core of the hot streak has small turning angle and keeping away from the vane wall.

In general, after passing the vane passage, the shape, the temperature value, and the radial and circumferential positions of the hot streak will change. The different variations of heat load and aerodynamic parameters of the rotor blade will mainly arise from the above variations.

Stator aerodynamic parameter variation

Compared with uniform total temperature inlet condition, hot streak has little effect on vane surface pressure distribu-tions, as shown in Fig.13. Vane surface pressure distributions of all hot streak schemes agree with that of uniform total tem-perature inlet condition at 10% and 65% vane height, only with a little difference at the rearward section of vane suction sur-face. In a subsonic turbine stage, rotor-stator interaction will lead to vane surface pressure fluctuation periodically, but the In addition, the inlet hot streak and its circumferential po-sition variations have little effects on the total mass-averaged total pressure, total temperature, and flow angle of the whole stator exit, compared to that of the uniform total temperature inlet condition, because the two kinds of inlet conditions have the same inlet mass-averaged total temperature and total pres-sure. But the hot streak will influence the aerodynamic parame-ter distribution at stator exit both in the circumferential and radial direction because of the temperature gradients. On ac-count of the rotor blade rotation in the circumferential direc-tion, the pressure and heat load fluctuations of rotor blade mainly arises from the stator exit circumferential parameter

variations, consequently, the emphasis is put on the circumfe-

rential parameter variations caused by the hot streak.

(a) Static temperature contours

(b) Static pressure contours

(c) Total pressure contours

Fig.14 Stator exit flow field under uniform total temperature

inlet condition (N=0)

It is necessary to illuminate the circumferential parameter variations in particular temperature and pressure at stator exit under the uniform total temperature inlet condition, as the base line for comparison. Fig.14 shows that the circumferential static temperature, total temperature and total pressure at stator exit are obviously non-uniform even under the uniform total tem-perature inlet condition. The variation of circumferential tem-perature and pressure comes from the different expansion level when the gases flow through the vane passage which is the in-herent characteristic of the cascade flow, as shown in Fig.14a and Fig.14b, and comes from the passage vortices as the black arrows indicated in Fig.14a. The passage vortices will lead to non-uniform temperature at exit plane, although the size of vor-tices is small. It is known that the passage vortex forms by the two branches of the horse vortex due to the interaction of the wall boundary layer and the leading edge. The fluid in boun-dary layer usually has low temperature in cooled turbine, so the passage vortices will lead to low temperature region on stator exit plane. Also, the passage vortices will lead to total pressure non-uniform at stator exit plane, as shown in Fig.14c. Com-monly, the passage vortices are the main sources of stator loss, so the passage vortices will influence the heat transfer of the downstream rotor blade not only by the low temperature but also by the flow characteristic. In addition, the wake should have an influence on the downstream flow and heat transfer. In general, the circumferential variation of temperature and pres-sure at stator exit under uniform total temperature inlet condi-tion should be consistent for the two vane passages.

Figure 15 shows the circumferential parameter variation at 65% vane height under hot streak and uniform total temperature

inlet conditions. Although the highest temperature region is slightly different in the radial direction among the six hot streak schemes, the 65% vane height still is a typical position for trend comparison. Among these parameters, flow angle, Mach num-ber, total pressure have little difference among the seven schemes, whereas total temperature, static pressure and velocity have great difference. Total temperature variation combines static temperature and velocity difference, total pressure differ-ence combines static pressure and velocity difference. Because total pressure difference among the schemes is little, the static pressure difference from one scheme to another comes from the velocity difference. Total temperature variation trend (Fig.15e) is not according with the static temperature variation trend (Fig.11) fully, also due to the velocity variation (Fig.15f)

Fig.15 Exit (x/C=103%) circumferential parameter distribu-

tion at 65% vane height (N=10)

The velocity difference at stator exit among the schemes

arises from the inlet hot streak. For the present simulation, out-

let averaged static pressure and inlet uniform total pressure is

specified. The high total temperature region due to the hot

streak at stage inlet has high total enthalpy, which corresponds

to the high flow velocity at stator inlet under the same exit

pressure. In other words, the circumferential velocity difference

among the schemes at stator exit is caused by the relative pitch

position variation of the hot streak at stator inlet.

Rotor aerodynamic parameter variation

The circumferential parameter variation at stator exit is the

main factor affecting the aerodynamic parameter variation of

the rotor, in particular the total pressure, total temperature, and

velocity. The analysis on the stator indicates that the hot streak

influences the circumferential parameter distribution and has

little effect on wall pressure distribution, because there is only

total temperature distortion at the stator inlet. For the rotor,

whereas, there are both inlet total temperature distortion and

inlet total pressure distortion as shown in Fig.16. Thus, the

aerodynamic parameter variation in the rotor should be greater

than that in the stator.

The analysis on the aerodynamic parameters focuses on

the surface pressure transient variation. The surface pressure

variation causes the flow change in blade passage, another as-

pect, the periodic variation of the surface pressure results the

transient aerodynamic force of the rotor blade fluctuating pe-

riodically, which is remarkable for vibration of the rotor blade.

Figure 17 shows that the surface pressure transient varia-

tion under hot streak inlet condition is obvious greater than that

under the uniform total temperature inlet condition. The tran-

sient variation of the surface pressure under uniform total tem-

perature inlet condition comes from the non-uniform circumfe-

rential pressure at stator exit as shown in Fig.14. Figure 18

gives the surface pressure transient variation within one rotor

domain period at three typical points of 65% blade height. The

point positions are shown in Fig.7a. At the blade leading edge

(Point 1), the fluctuation range of pressure is maximal, the

maximal transient pressure difference is about 10%p0*. The

fluctuation range increases by the hot streak but the value is

small relatively. At the suction surface (Point 2), the pressure

fluctuation range is smaller than the leading edge as a whole,

and the effect of the hot streak is relatively greater than leading

edge. Also, the pressure fluctuation on the suction surface is

very sensitive to hot streak. When the rotor blade passes the

vane exit where hot streak effect exists (such as N=6-26), the

fluctuation range increases obviously. At the pressure surface

(Point 3), the pressure fluctuation is very small, i.e., only about

2%p0* transient pressure difference under uniform total tem-

perature inlet condition, but the hot streak increases the tran-

sient pressure difference to about 4%p0*.

Fig.16 Total temperature and total pressure contours in

Fig.17 Surface pressure distributions at 65% blade height

(N=0-39)

height of three typical points

Because the circumferential parameters are non-uniform in the two vane passages for hot streak inlet condition, the fluctua-tion periodicity of the surface pressure altered from 9 degree to 18 degree circumferential interval, compared with the uniform total temperature inlet condition.

Vane wall heat flux variation

The effect of the hot streak on heat load is more important than on aerodynamic parameters in gas turbine from the view point of reliability. Firstly, hot streaks lead to heat load non-uniform in radial direction whichever for vane 1 or vane 2 be-cause of the radial temperature gradient. Secondly, hot streaks lead to non-uniform heat load for the two vanes in one hot streak period because of the circumferential temperature gra-dient. The heat load distribution caused by the relative position between the hot streak and the vane is very important for vane cooling design.

Figure 19 shows that at 65% vane height, q/q0 of vane 1 decreases with the circumferential turning of hot streak from RPLE=-2% to 81%. The vane 1 has the maximum q/q0 when RPLE=15%, at the same time, vane 2 has the minimum q/q0. Hence, RPLE=15% causes the maximum heat load non-uniformity between vane 1 and vane 2 in one hot streak period. RPLE=-2% has the similar behavior compared to RPLE=15%, because the highest total temperature region of hot streak im-pinges upon the vane 1 leading edge, too. For RPLE=65%, vane 1 has approximately the same q/q0 as uniform total tem-perature inlet condition. Also, vane 1 and vane 2 has approx-imate q/q0. For RPLE=81%, q/q0 of vane 1 is lower than uni-form total temperature inlet condition. Vane 2 has the opposite behavior compared with vane 1 when the pitch position of hot streak changed. Figure 20 shows that at the relative low tem-perature region, at 10% vane height, the q/q0 is smaller than at 65% vane height for all hot streak schemes. The circumferential position variation of the hot streak has little effects on q/q0

The analysis above indicates that RPLE=15% causes the maximal circumferential and radial non-uniformity in heat load. RPLE=65% can obtain almost the same heat load on vane 1 and vane 2, and the minimal non-uniformity in heat load in the radial direction of the vane.

Blade wall heat flux variation

Thermal fatigue of the rotor blade is one important aspect for reliability. The thermal fatigue comes from temperature fluctuation on the rotor blade surface. The hot streak is one important factor that can cause the temperature fluctuation on

the rotor blade due to its circumferential temperature gradient. Although it is not very clear how much the effect can be caused to the blade fatigue life, it is necessary to give the transient

Fig.21 q/q0 distributions of rotor blade (N=0-39)

The periodic variation of q/q0 on the rotor blade is mainly

caused by the exit circumferential temperature variation of the

stator. Wake flow and passage vortices also have some effects

too, but the effects are relative weak. Figure 21 gives q/q0 com-

parisons between uniform total temperature inlet condition and

hot streak inlet condition (for example RPLE=48%). It indi-

cates that q/q0 fluctuates in a small range under uniform total

temperature inlet condition, whereas, hot streak inlet condition

leads to q/q0 fluctuating dramatically, in particular at the blade

leading edge, forward portion of the suction surface, and rear-

ward portion of the pressure surface. Moreover, the fluctuation

range on suction surface is greater than on pressure surface. At

65% blade height is greater than at 10% blade height.

The behaviors of the other hot streak schemes are similar

to RPLE=48%. This result indicates that hot streaks increase

the fluctuation range of wall heat flux and the non-uniform de-

gree of wall heat load.

Figure 22 describes the transient variation process of q/q0

on the blade suction surface. The process pulls together many

factors including temperature field variation caused by hot

streak, 3D viscous flow in stator vane passage such as passage

vortex and wake flow, and 3D viscous flow in rotor blade pas-

sage itself. Hot streak influences heat transfer mainly on for-

ward section of the blade passage especially suction surface,

where the influence through the whole span. At rearward sec-

tion of the rotor blade passage, 3D separation line on blade suc-

tion surface limits the transports of hot streak. The q/q0 con-

tours at N=28-36 shows the effect of the passage vortices

coming from the stator passage, as the black arrows denoted,

this effect only exists in local region. In addition, the effect of

hot streak and passage vortex appear at different time step,

which indicates the circumferential position difference between

the hot streak and passage vortex at stator exit. The effect of

vane wake is not showed distinctly because of the weak effect.

N=0 N=4

N=8 N=12 N=16 N=20

N=24 N=28 N=32 N=36

Fig.22 Transient variation of q/q0 on blade suction surface

(RPLE=48%)

Figure 23 shows transient temperature contours at 65%

blade height. It can be seen that rotor blades experience circum-

ferentially non-uniform temperature field, and for the same

blade, there are phase differences between the highest and low-

est temperature among hot streak schemes. For the hot streak

flow in blade passages, the typical V-shape temperature distri-

bution agrees with the results in Dorney et al [7], also the hot

streak fluid moves towards the blade pressure surface distinctly

under the similar inlet condition RPLE=48%. Dorney et al [7] focuses on the time-averaged pressure surface temperature “hot spot”. The present study pays more attention to the transient variation of heat load on both the suction surface and pressure surface. Figure 24 shows the transient periodic fluctuation of heat flux at three typical points, the point positions are shown in Fig.7a. As the pressure variation trend, the leading edge (Point 1) has the maximal transient q/q0 difference among the schemes and the pressure surface (Point 3) has the minimal transient q/q0 difference. Unlike the pressure fluctuation, two vane passages have one q/q0 peak value corresponding to one peak hot streak temperature. In addition, the phase difference

relates to the vane exit temperature filed directly.

RPLE=-2% RPLE=15% RPLE=31%

RPLE=48% RPLE=65% RPLE=81%

Fig.23 Transient temperature contours at 65% blade height

surface

The transient variation of the averaged q/q0 exhibits the

combined unsteady effects of the hot streak, as shown in

Fig.25. This variation can represent the total heat load that the

rotor blade endures at each time step. In general, the fluctuation

range of the averaged q/q0 is very small under the uniform total

temperature inlet condition, as the dash line denoted, but hot

streak increases the fluctuation range of heat load greatly. The

fluctuation range of the averaged q/q0 on the pressure surface

shows the difference about the hot fluid accumulating on the

pressure surface when the hot streak circumferential position

varying. RPLE=31% has the highest peak averaged q/q0, and

RPLE=-2% has the smallest peak averaged q/q0. This result

confirms the trend analysis in He et al [6] basically, i.e., the hot

fluid obviously accumulates on the pressure surface when the

hot streak locating at the middle of the vane passage inlet, and

it is not easily identifiable when the hot streak impinging the

vane leading edge. The fluctuation range of the averaged q/q0

on the suction surface is greater than on the pressure surface for

all hot streak schemes. The difference among the hot streak

schemes exists in maximum and minimum value of the aver-

aged q/q0 and the time of the extreme value appearing, and the

value difference is more important. Among all hot streak

schemes, RPLE=31% has the highest averaged q/q0 at the ex-

treme point just like the behavior of the pressure surface, and RPLE=81% has the lowest averaged q/q0 at the extreme point. RPLE=65% and RPLE=-2% have similar behavior to RPLE=81% and the three hot streak inlet conditions have a common point that the highest temperature regions are close to the vane suction surface at stage inlet. Considering the heat load of the vane and blade together, RPLE=65% has the mi-nimal non-uniform degree of heat load on the vanes, also re-sults in the small averaged heat load fluctuation on the rotor blade.

CONCLUSIONS

The circumferential position variations of hot streak at stage inlet lead to the variations of position, shape and value of high temperature region at vane exit. After the migrations through the vane passages, the circumferential interval of the highest temperature region becomes wide and the radial posi-tion becomes high gradually with the hot streak turning from TE1 pressure side to TE2 suction side. At vane exit, if the high temperature region is close to the pressure surface, then it spreads in the radial direction greatly. And if the high tempera-ture region closes to the suction surface, then it spreads in the circumferential direction greatly. RPLE=31% has the highest peak temperature.

Compared with uniform total temperature inlet condition, hot streak has little effects on the vane surface pressure and stator exit total mass-averaged parameters, regardless of the hot streak circumferential position variation. But the hot streak has great effects on circumferential distributions of the aerodynam-ic parameters, especially on total temperature, static pressure, and velocity. The velocity variation arises from the stage inlet total temperature distribution, and the relatively high total tem-perature region results in the relatively high velocity at stage inlet.

Under hot streak inlet conditions, two vanes facing one hot streak have different wall heat load obviously. RPLE=15% has maximal non-uniform level of vane heat load in the circumfe-rential and radial direction due to impinging upon the leading edge of vane 1. RPLE=65% has minimal non-uniform level of vanes heat load, which is close to that of uniform total tempera-ture inlet condition.

Surface pressure and heat load of rotor blade fluctuates pe-riodically with time even under uniform total temperature inlet condition, but hot streaks increase the fluctuation range greatly especially heat load. For the surface pressure, the maximal ef-fect of hot streak occurs on the pressure surface, the fluctuation range increases about 100%. For the heat load, the effect lie on the relative pitch position of the hot streak, RPLE=31% has the largest range fluctuation of the averaged heat load, RPLE=81% and 65% have relatively small fluctuation range of the averaged heat load. In addition, the hot streak alters the fluctuation pe-riod of surface pressure and heat load from 9 degree to 18 de-gree.

In general, the hot streak RPLE=65% obtains similar heat load for the two vanes corresponding to one hot streak, at the same time, obtains small fluctuation of the averaged heat load on rotor blade.

ACKNOWLEDGEMENTS

The authors wish to acknowledge the financial support from the National 973 Research Program of China. REFERENCES

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组件热斑效应

组件热斑效应 众所周知为了使组件达到最高的功率输出,光伏组件中的单体电池须具有相似的特性,对于组串及阵列也是如此。但在实际使用过程中,可能出现电池裂纹或不匹配、内部连接失效、局部被遮光或弄脏等情况,导致一个或一组电池的特性与整体不谐调。失谐电池不但对组件输出没有贡献,而且会消耗其他电池产生的能量,导致局部过热。这种现象称为热斑效应。当组件被短路时,内部功率消耗最大,热斑效应也最严重。 热斑效应不仅会严重影响组件的性能和使用寿命,还有可能引发燃烧及火灾,给电站带来财产损失和人员伤害,因此有效的判断热斑效应的发生及严重性是电站长期的工作。下左图是电站现场发生的组件背板灼烧现象。 对于热斑效应的判断,切记勿用手去触摸组件,因为当热斑发生时,组件的局部温度非常高,极有可能造成灼伤。运维人员应选择相应的测试仪器去对组件整体温度进行测试判断,并提早发现组件是否已经存在局部温度异常。此时选用最方便最快捷的测试仪器即是红外热像仪。红外热像仪可以全方位拍摄整个组件甚至阵列的温度分布情况,及时发现热斑所在。并通过软件全面了解组件当前的发热情况,对于明显有热斑的组件可以清楚判断,同时可对组件中尚不明显的热点进行分析判断。如上右图所示。 从图中可看出组件靠近地面的部位均存在一定程度的热斑效应,这是热斑效应发生概率较高的部位,原因是:(1)这部分组件最容易被遮挡,被遮挡的时间也最长;(2)灰尘覆盖最严重,有时候清洗的不干净时,这部分囤积的灰尘也越多。(3)靠近地面,通风较差,散热不佳。因此发生热斑效应的概率较高。当然引起热斑效应的原因并不止这些,组件本身的性能差别,是否存在隐裂,是否有损伤等等也会造成热斑效应。 HT测试仪器建议在运维过程中,对于已经存在热斑效应的组件,需要对其进行I-V曲线测试判断其功率下降的比例,对于热斑效应较严重的组件可考虑更换组件,避免对整个组串造成过大影响。对于尚未存在热斑效应的组件,最好进行抽查,对部分组件的I-V曲线进行测试,这样可以提前发现造成组件功率下降的原因,并及时改进。

热载流子效应对器件可靠性的影响

重庆邮电大学研究生堂下考试答卷 2011-2012学年第2学期考试科目微电子器件可靠性 姓名徐辉 年级2011级 专业微电子与固体电子学 学号S110403010 201 20122年5月25日

热载流子效应对器件可靠性的影响 徐辉 (重庆邮电大学光电工程学院,重庆400065) 摘要:介绍了几种热载流子以及MOSFET的热载流子注入效应。在此基础上总结了热载流子注入效应对MOS器件可靠性的影响。随着MOS器件尺寸的缩小和集成电路规模的增大,热载流子效应显得更加显著。最后介绍了几种提高抗热载流子效应的措施。 关键词:热载流子;热载流子注入效应;可靠性 Effects of Hot-carriers Injection Effect on the Reliability Xu Hui (College of Photoelectric Engineering,Chongqing University of Posts and Telecommunications,Chongqing, 400065,P.R.China) Abstract:The effect of hot carrier and the MOSFET hot-carriers injection are reviewed.On this basis,the hot-carriers injection effect on the reliability of MOS devices are summed up.With the increasing size of MOS devices shrink in size and integrated circuits,the hot-carriers effect is even more significant.Finally,several measures to improve the thermal carrier effects are introducted. Key wards:hot carrier;hot-carriers injection effect;reliability 0前言 随着VLSI集成度的日益提高,MOS器件尺寸不断缩小至亚微米乃至深亚微米级,热载流子效应已成为影响器件可靠性的重要因素之一。从第一次意识到热载流子可导致器件退化以来,有关MOSFET热载流子效应的研究已持续了近30年。热载流子注入效应对MOS器件性能的影响也越来越引起人们的关注。 1热载流子 当载流子从外界获得了很大能量时,即可成为热载流子。例如在强电场作用下,载流子沿着电场方向不断漂移,不断加速,即可获得很大的动能,从而可成为热载流子。对于半导体器件,当器件的特征尺寸很小时,即使在不很高的电压下,也可产生很强的电场,从而易于导致出现热载流子。因此,在小尺寸器件以及大规模集成电路中,容易出现热载流子。 在使用条件下,MOSFET会遇到四种类型的热载流子[1]; 沟道热载流子(CHC);衬底热载流子(SHC),漏端雪崩热载流子(DAHC);和二次产生热电子(SGHE)。 沟道热载流子(CHC):热电子来源于表面沟道电流,是从源区向漏区运动的电子,在漏结附近受到势垒区电场加速,电子获得了能量而被加速,成为热电子。 衬底热载流子(SHC):热电子来源于衬底电流,在势垒区电场的加速下运动到Si-SiO2界面,其中部分电子的能量可以达到或超过Si-SiO2势垒高度,便注入到栅氧化层中去。 漏端雪崩热载流子(DAHC):晶体管处在饱和状态时,一部分载流子在夹断区域与晶格原子相撞,通过碰撞电离,激发电子-空穴对。

航空发动机涡轮叶片

摘要 摘要 本论文着重论述了涡轮叶片的故障分析。首先引见了涡轮叶片的一些根本常识;对涡轮叶片的结构特点和工作特点进行了详尽的论述,为进一步分析涡轮叶片故障做铺垫。接着对涡轮叶片的系统故障与故障形式作了阐明,涡轮叶片的故障形式主要分为裂纹故障和折断两大类,通过图表的形式来阐述观点和得出结论;然后罗列出了一些实例(某型发动机和涡轮工作叶片裂纹故障、涡轮工作叶片折断故障)对叶片的故障作了详细剖析。最后通过分析和研究,举出了一些对故障的预防措施和排除故障的方法。 关键词:涡轮叶片论述,涡轮叶片故障及其故障类型,故障现象,故障原因,排除方法

ABSTRACT ABSTRACT This paper emphatically discusses the failure analysis of turbine blade.First introduced some basic knowledge of turbine blades;The structure characteristics and working characteristics of turbine blade were described in she wants,for the further analysis of turbine blade failure Then the failure and failure mode of turbine blades;Turbine blade failure form mainly divided into two major categories of crack fault and broken,Through the graph form to illustrate ideas and draw conclusions ;Then lists some examples(WJ5 swine and turbine engine blade crack fault,turbine blade folding section)has made the detailed analysis of the blade.Through the analysis and research,finally give the preventive measures for faults and troubleshooting methods. Key words: The turbine blades is discussed,turbine blade fault and failure type,The fault phenomenon,fault caus,Elimination method

组件热斑效应的原因与防护

组件热斑效应原因和运维防护措施 曹晓宁1 闻震利2 吴达 1 ( 1. 中广核太阳能开发有限公司 100048; 2. 镇江大全太阳能有限公司 212211) 摘要:光伏电站中组件在运行中存在很多因素引起功率损耗并可能导致安全问题,热斑效应会造成组件功率的大幅度下降,而且是比较严重的安全隐患。在组件生产过程、现场施工和运行维护中可以对技术指标提出要求或采取相应的措施来防护热斑效应。为了减少运维工作量,提供效率,监控系统可以对组件的电流和电压进行监测并进行逻辑判断,可帮助运维人员进行针对性的排查,提高光伏电站运行的安全可靠性。 光伏发电是人类解决能源危机和环境问题的必由之路,在过去的二十年里光伏发电产业有了迅猛的发展,权威能源机构预测在本世纪中叶光伏发电会能为人类主要的供电方式之一。太阳电池组件是光伏电站的核心元件,组件的性能和安全可靠性直接决定了光伏电站的运行效率。目前组件的标称功率是在标准测试环境下(标准条件具体是指:温度25℃,光谱分布AM1.5,辐照强度是1000W/m 2)的发电功率,而在实际运行环境中,由于温度、辐照强度、光谱失配等因素会影响组件的实际发电功率。在实际应用中,组件的阴影遮蔽是不可避免的问题,阴影遮蔽会造成功率损失,而且会导致局部发热,产生安全隐患,即热斑效应。本文对热斑的成因和热斑效应的防护措施进行探讨。 1、热斑效应 晶硅组件是由多个太阳电池片串联组成,当串联支路中的一个太阳电池被遮挡时,将被当作负载消耗其他的太阳电池所产生的能量,被遮蔽的太阳电池此时会严重发热,称为热斑效应,如图1所示。热斑效应会严重影响组件的输出功率,同时会破坏太阳电池的性能。有光照的太阳电池所产生的部分能量,都可能被遮蔽的电池所消耗,热斑效应时组件温度分布如图2所示,可以看到被遮挡电池的温度明显高于其它电池。 图 1 热斑效应原理示意图 图2 热板效应时组件的温度分布图 2、热斑效应的防护措施 组件中电池片的电流失配、电池片破损、组件虚焊和污损遮挡等原因都会引起电池发热,为了防止热斑效应对光伏电站造成发电量损失及对太阳电池造成损伤,应该在组件生产、现场施工和运行维护过程中采取相应的措施来减少热斑效应发生的风险,降低其危害。 2.1组件生产过程控制 首先对太阳电池进行电流分档,减少组件中串联太阳电池之间的电流失配,另外对组件进行功率分档后,仍要进行电流分档;其次在电池两端并联旁路二极管,即在组件中安装旁路二极管;再次对太阳电池的反向漏电进行控制,太阳电池承受12V 反向电压热斑 热斑

第一次综述热载流子注入效应对MOS器件性能的影响讲解

热载流子效应及其对器件特性的影响 组长:尹海滨09023105 整合资料撰写综述 组员:马祥晖09023106 查找问题三资料 王小果09023128 查找问题二资料 李洋09023318 查找问题一资料

目录 一绪论————————————————————————————————3 二正文主题——————————————————————————————4 1热载流子与热载流子注入效应—————————————————————4 1.1载流子的概念 1. 2热载流子的概念及产生 1. 3热载流子注入效应 1.4热载流子效应的机理 2热载流子注入效应对MOS器件性能的影响———————————————6 2.1热载流子对器件寿命的影响 2. 2热载流子效应的失效现象 2.2.1雪崩倍增效应 2.2.2阈值电压漂移 2.2.3 MOSFET性能的退化 2.2.4寄生晶体管效应 2.3热载流子注入对MOS结构C-V和I-V特性的影响 2.3.1热载流子注入对MOS结构C-V特性的影响 2.3.2热载流子注入对MOS结构I-V特性的影响 3提高抗热载流子效应的措施——————————————————————10 3.1影响热载流子效应的主要因素 3.2提高抗热载流子效应的措施 三结论————————————————————————————————12 四主要参考文献————————————————————————————12

一绪论 随着科学技术的发展,半导体器件在未来将会有着良好的发展前景,据世界半导体贸易统计歇会(WSTS)日前发布的一份预测报告,世界半导体市场发展未来三年将会保持两位数的增长,这份报告中还表明,全球半导体业之所以能保持高增长,集成电路IC芯片的高需求功不可没,给全球半导体业注入了新的活力。在最近三年里,三网融合的大趋势有力的推动着芯片业的发展。无论是在移动通信业,无线数据传输业,还是PC机芯片都有着良好的发展趋势。而缩小芯片体积和提高芯片性能是阻碍集成电路发展的两大重要因素,为了进一步缩小芯片体积,科学家们正在研制一系列的采用非硅材料制造的芯片,例如砷化镓,氮化镓等;另外芯片器件性能的提高也是重中之重,其中芯片器件可靠性是衡量其性能的重要指标,尤其是在航天,航海等军事方面尤为重要。 本综述报告讨论的就是对器件特性和可靠性的影响因素之一的热载流子效应及其应用。

光伏电站理论发电量计算及影响因素

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辐照度为800W/m2,电池温度20℃时,250W的组件输出功率只有183W,为标况下的73.2%。 2、效率 理论上,尺寸、标称功率相同的组件,效率肯定是相同的。光伏组件是由电池片组成,一块光伏组件通常由60片(6×10或72片(6×10电池片组成,面积分别为 1.638 m2(0.992m×1.652m和 3.895 m2(0.992m×1.956m。 辐照度为1000W/m2时,1.638 m2组件上接收的功率为1638W,当输出为250W时,效率为15.3%,255W时为15.6%。 3、电压与温度系数 电压分开路电压和MPPT电压,温度系数分电压温度系数和功率温度系数。在进行串并联方案设计时,要用开路电压、工作电压、温度系数、当地极端温度(最好是昼间进行最大开路电压和MPPT电压范围的计算,与逆变器进行匹配。 二、影响光伏组件的两个效应 1、热斑效应 一串联支路中被遮蔽的太阳电池组件,将被当作负载消耗其他有光照的太阳电池组件所产生的能量,被遮蔽的太阳电池组件此时会发热,这就是热斑效应。

热斑

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凡是串连就会由于组件的电流差异造成电流损失; 凡是并连就会由于组件的电压差异造成电压损失; 组合损失可以达到8%以上,中国工程建设标准化协会标准规定小于10%。 注意: (1) 为了减少组合损失,应该在电站安装前严格挑选电流一致的组件串联。 (2) 组件的衰减特性尽可能一致。根据国家标准GB/T--9535规定,太阳电池组件的最大输出功率在规定条件下试验后检测,其衰减不得超过8%。 (3) 隔离二极管有时候是必要的。 5、温度特性 温度上升1℃,晶体硅太阳电池:最大输出功率下降0.04%,开路电压下降0.04%(-2mv/℃),短路电流上升0.04%。为了避免温度对发电量的影响,应该保持组件良好的通风条件。 6、灰尘损失 电站的灰尘损失可能达到6%!组件需要经常擦拭。 7、MPPT跟踪 最大输出功率跟踪(MPPT)从太阳电池应用角度上看,所谓应用,就是对太阳电池最大输出功率点的跟踪。并网系统的MPPT功能在逆变器里面完成。 8、线路损失

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