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GTX 100 Turbine Section Measurement Using
a Temperature Sensitive Crystal Technique.
A Comparison With 3D Thermal and
Aerodynamic Analyses
Mats Annerfeldt
Sergey Shukin
Mats Björkman
Agne Karlsson
Anders Jönsson
Elena Svistounova
Demag Delaval Industrial Turbomachinery AB
Finspong, Sweden
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GTX 100 TURBINE SECTION MEASUREMENT USING A TEMPERATURE
SENSITIVE CRYSTAL TECHNIQUE. A COMPARISON WITH 3D THERMAL AND
AERODYNAMIC ANALYSES
Mats Annerfeldt, Sergey Shukin, Mats Björkman, Agne Karlsson, Anders Jönsson
Elena Svistounova
Demag Delaval Industrial Turbomachinery AB, Finspong, Sweden
1. Abstract
In modern gas turbine engines, higher and higher turbine inlet temperature is used in order to
increase the efficiency. To achieve a high benefit from increased temperature level it is
necessary to minimise the amount of cooling air, which reduces the thermal cycle efficiency.
The difficulty in turbine design is to find the optimal path to increase the efficiency without
sacrificing the component lifetimes.
Modern gas turbine materials usually suffer a steep decrease in material properties when a
certain temperature is exceeded. It is extremely important to know the component
temperatures in real engine conditions with good accuracy, in order to be able to predict the
component lifetimes.
For the heavily cooled components, the main damage mechanism is often thermo-mechanical
fatigue, TMF, caused by the thermal gradients within the component. With more traditional
instrumentation using thermocouples it is not possible to install enough measuring points on
the component to really catch the gradients. Thermal paints show the gradients, but the
commercially available paints are too sparse between the temperature transitions and are often
hard to evaluate with the necessary accuracy of temperature level. The Thermo-crystal
method enables measurement of the temperature with good accuracy in many points on the
same component.
This paper presents the way in which such a measurement was performed under real engine
conditions and shows some of the results. Both gas and metal temperatures for stationary
components as well as rotating blades were measured with Thermo–crystals during the same
test run.
Furthermore, the results from the measurement are compared to the calculated temperature
field of the same component using a 3D heat transfer conjugate model, from which the
temperature field used for lifetime predictions is taken.
The gas temperatures are used for comparing and tuning of the 3D multistage CFD model
used to calculate the temperature boundary conditions for the thermal model of the
component. A comparison between measured and calculated temperature attenuation is
presented in the paper.
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2. Nomenclature
T
T*
Tw*
τ
2*θ
temperature [C]
Stagnation temperature absolute frame of reference [C]
Stagnation temperature in relative frame of reference[C]
time [min]
Diffraction angle
Mea = measurements
S3D = MBStage3D
3. Introduction
Gas turbine plant owners are, for obvious reasons, very interested in keeping the intervals
between overhauls as long as possible. On the other hand, a forced outage due to component
failure or premature exchange outside the planned inspections should be avoided. It is
therefore of greatest importance to the gas turbine manufacturers to be able to predict
component lifetimes with good accuracy. A detailed knowledge of the temperatures being
exposed to different components during operation is then necessary.
It is very difficult to predict the temperatures with necessary accuracy in all positions of all
components only by using calculation methods. Temperature measurements are needed to
complement the calculations in order to reach a high level of confidence in life expectancy.
The measurements also provide the possibility to detect any life issues at an early stage, or to
identify potentials to reduce the cooling air consumption, improving the overall engine
performance.
The GTX100, a 45 MW industrial gas turbine with 37% efficiency, has successfully
accumulated more than 110 000 operation hours. A number of component upgrades have been
introduced since the original launch and a new fingerprint of the complete turbine section was
taken during a comprehensive measurement in 2003. The instrumentation used in this test
included more than 2 300 measuring points, complemented also with thermal paint. A total
number of 1 975 thermo-crystals, 237 thermocouples and 110 pressure taps were used for the
test of the 3-stage turbine. This paper will focus on the thermo-crystal technique, which gives
an excellent mapping of the temperature distribution in turbine vanes and blades.
The evaluation process for the measurement is shown in Fig 1. In the following, the results for
the temperature attenuation throughout the turbine stages and the thermal results for blade 2
will be presented.
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Measured Gas
temperatures and
pressures in the
flow path
Measured metal
temperatures
3D aero model
of all 3 turbine
stages
Refined 3D aero
local model for each
component
Mechanic Integrity
model
3D conjugate heat
transfer model
Predicted life of the component Hours,
Cycles
Metal temperature
distribution
Free stream temperature,
pressure and velocity
.
distribution
Boundary condition at inlet and
outlet of the component
Figure 1 Evaluation process schematic picture
4. The experimental method
When the material of the JMTK (Russian abbreviation for the maximum temperature crystal
measurement) is exposed to neutron radiation, its atomic lattice will be shifted so that the
distance between the atomic planes will be larger. The distance between atomic planes can be
measured using an X-ray diffraction microscope. When a JMTK, which has been illuminated,
is heated up, the distortion induced by the neutron rays will relaxate. The relaxation is
dependent on time and temperature. A higher temperature will make the relaxation go faster.
If the diffraction angle 2*θ is measured after the crystal has been heated up, and the time
during which it has been heated is known, it is possible to look up the temperature it has been
at from a calibration diagram. See Figure 2.
For exposure times over 10 minutes the dependence upon time is not so strong, as can be seen
from Fig 2. If the crystal goes through a transient cycle, which contains several different
temperature levels, the whole cycle has to be integrated. The maximal temperature during the
cycle will be dominating, but lower temperatures will also contribute to the relaxation. The
integration is made by calculating an equivalent hold time at maximal temperature of the
transient.
To be able to accurately calculate the equivalent hold time usually a few thermocouples are
installed on typical positions to record the temperature transient. The temperature transients
for three thermocouples on vane 2 are seen in Fig. 3 The hold time at maximal power output
is approximately 20 minutes.
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Evaluation
diagramdiagram
for Thermocrystals
Calibration
for JMTK of SiC
(Here shown for every 100 C only)
167
166
165
100
200
164
300
400
500
Angle [2* ]
163
600
700
162
800
900
161
1000
1100
1200
160
1300
1400
159
158
157
1
10
100
Equival ent Hold Time [min]
Figure 2 Typical Calibration diagram for JMTK crystals This diagram is batch specific,
and the figure above is only schematic
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900
850
800
750
700
650
600
550
500
450
400
350
300
250
200
150
100
50
0
90
85
80
75
70
65
60
55
50
45
40
35
30
25
20
15
10
5
0
TTVW 0240204
P_el [MW]
Temperature [C]
Vane 2 M etal temperatures during the test
TTVW 0240304
TTVW 0240404
P_el [MW ]
75
70
65
60
55
50
45
40
35
30
25
20
15
10
5
0
Time [min]
Figure 3. Transient Temperature response for a TC on Vane 2, aerofoil.
The accuracy of the method is claimed by the crystal supplier to be +/- 10C . This has been
verified by blind tests for some crystals which had been put in an oven at a well known
temperature for a known time. This experiment confirmed the claimed accuracy.
The measuring interval is from 200C to 1400C. For high temperatures the exposure time is
limited.
In this particular test crystals with a diameter of 0.2 mm were used
The metal temperatures were measured by installing the crystals in circular grooves with
diameter, d=0.5mm and depth 0.5mm, using a thermo-cement technique to glue the crystals
into the grooves.
Another advantage of the crystals is that the surface is completely smooth after the
installation, as distinct from many thermocouple installations, and gives better accuracy than a
thermal paint test. It is also possible to install many crystals on a single component. It is an
advantage be able to measure the thermal gradients from a single component instead of doing
a jigsaw puzzle from temperatures measured at several components, as is usually necessary
when measuring with thermocouples. To actually measure the correct gradients is essential
when the main damage mechanism is TMF, as is usual, particularly for first stage components
of intensively cooled components.
The yield rate of the metal temperature crystals during this test is quite good. 95% of the 1975
installed crystals remained in the grooves throughout the test and gave an evaluated
temperature.
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The crystal technique can also be used to measure the gas temperature profile entering the
vane or blade. In this case, crystals are installed at the tip of small ceramic pins attached to the
leading edge. The high heat transfer coefficients in the gas channel environment provide
enough gradient along the pin so that the pin will be insensitive to the conduction of heat from
the pin to the blade, which is colder than the gas. The difference between measured
temperature and real temperature is only about 4C as a maximum in the presented results
(Blade 1). Figure 4 below shows the installation on blade 2.
Figure 4 Blade 2 instrumented with crystals measuring Gas temperature Tw*
In the tests performed on GTX100 reported here the inlet gas temperature of all 6 rows of the
3 stage turbine were instrumented with gas temperature measurement. The yield for the rotor
blades gas temperature measurement was unfortunately less than for the metal crystals and the
vane gas temperatures, only about 80% out of 120 installed crystals.
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5. Measured gas temperature attenuation vs. gas temperatures from 3D NS multistage
calculations
When designing the cooling system for cooled components in modern gas turbines it is
necessary to spend a minimum of cooling air, in order to minimise the negative effect on the
efficiency from the cooling air injection. This means the cooling has to be tailor-made for the
gas temperature distributions of the particular turbine.
Most CFD codes have difficulty in correctly predicting the mixing of the flow as it passes
through the turbine, and thus the temperature attenuation through it. Usually the average
temperature can be predicted with good accuracy, but it is more difficult to predict the shape
of the temperature profile. In this paper the measurement results are compared to the results
from a 3D NS calculation using the code Stage3D.
This kind of test also provides valuable information for verification of CFD codes.
5.1. 3D Description of the CFD calculations
This chapter give a brief description of the CFD calculations performed in order to have
boundary conditions for the 3D conjugate heat transfer code. A schematic picture of the
evaluation process was shown in the introduction, see Fig.1.
An in-house 3D Navier-Stokes solver, MBStage3D, was used to calculate the gas temperature
and other necessary boundary conditions for the conjugate heat transfer calculations. The
CDF calculations were divided into two steps. First a regular simulation of the whole turbine
was performed, then each component was simulated in a separate model, in order to increase
the accuracy by tuning each model against the measured radial temperature distribution.
All simulations were steady, the Spalart – Allmaras one-equation turbulence model with wall
functions was applied and each component was modelled using straight H-mesh grids
containing 350 – 600 thousand nodes (2.5 million in the full turbine model), resulting in 10 <
y+ < 300. In the full turbine model, the inlet boundary conditions (stagnation temperature and
pressure, flow angles and turbulent properties) were taken from a combustor CFD calculation,
the cooling air boundary conditions from a Secondary Air Flow (SAF) calculation and for the
outlet static pressure measured values were used. Inlet and outlet boundary conditions for the
separate models were taken from the full turbine simulation, but the inlet stagnation
temperature was adjusted according to the measurements. Additionally the cooling air
boundary conditions were enhanced in the separate models; these were tuned against previous
field experience and the measured metal temperatures, which indicate how the cooling air
behaves. Fig. 5a shows the computational domain of the full turbine model, one passage in
each component was modelled with assumed periodicity and averaging mixing planes.
Additionally the predicted relative free stream stagnation temperature close to the metal
surface of blade 2, is shown in Fig. 5b. These patterns agree well with traces on blades that
have been under operation in a real machine.
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Turbine simulation
followed by separate
models with enhanced
boundary conditions
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Figure 5a Computational domain of the
full turbine model
Trace of
trailing edge
cooling air
ejection
Boundary conditions
for 3D conjugate heat
transfer calculations
View of pressure side
View of suction side
Cold streaks
following the
secondary
flow up/down
on the aerofoil
Figure 5b. Relative free stream
stagnation temperature calculated using
separate model of blade 2. ∆T*w = 10K
Cold purge air coming out
from the cavity between
Vane 2 and Blade 2 up on
the platform surface
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5.2. Measured gas temperature attenuation vs gas temperatures from 3D NS multistage
calculations
In order to see how well the full turbine simulation predicts the mixing of cooling air to the
main stream, the calculated and measured stagnation temperatures were compared. Fig.6
shows the results for all the components except the first stage. The measured values are
shown by symbols, e.g. 'Mea T*5' is the measured temperature in the second vane at tangential
position no. 5 and 'Mea T*av' is the average of 'Mea T*5', 'Mea T*22' and 'Mea T*40'. The black
solid lines show the result of the MBStage3D calculations.
The agreement between measurements and calculations is generally fair, the difference is less
than +/- 20K.
Generally the measured temperatures from the blades are less scattered between different
blades in the same stage than for the vanes. This is logical as the blades scan over the whole
turn, and therefore feel the average temperature. The vanes are stationary and will thus show
the temperature in the particular section where they are positioned, which can differ due to
tangential temperature variations from the combustor.
Regarding the shape of the calculated stagnation temperatures, the general trends are captured
but the calculation shows a more oscillating behaviour (the number of measurement points
should have been enough to capture this). This indicates that the mixing of the cooling air
with the hot main stream was too slow in the calculations.
In summary, it looks promising for future evaluations of other cases where measurements do
not exist. This comparison shows a quite good agreement between calculated and measured
temperatures.
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Blade 2 Radial Gas Tem perature distribution
Vane 2 Radial Gas Tem perature distribution
1
1
0,9
0,9
M ea: T*av
M ea: T*22
0,6
0,5
M ea: T*40
0,4
S3D Calc.
0,3
fraction of span [-]
fraction of span [-]
M ea: T*5
0,7
M ea: Tw*4
0,6
0,3
0,1
0
820
1100
Vane 3 Radial Gas Tem perature distribution
1
0,9
0,9
0,8
0,5
0,4
M ea:T*13
M ea:T*23
S3D Calc.
0,3
920
0,7
0,6
M ea: Tw*av
0,5
M ea: Tw*1
M ea: Tw*3
0,3
0,2
0,1
0,1
T* [C] dT=10 C 800
betw een gridlines
850
M ea: Tw*2
0,4
0,2
0
750
840 850 860 870 880 890 900 910
Tw * [C] dT= 10 C betw een gridlines
0,8
M ea: T*av
fraction of span [-]
0,6
830
Blade 3 Radial Gas Tem perature distribution
1
0,7
S3D calc.
0,4
0,2
T* [C]1000
dT=10 C betw een 1050
gridlines
M ea: Tw*7
0,5
0,1
0
950
M ea: Tw*1
0,7
0,2
fraction of span [-]
M ea: T*av
0,8
0,8
0
600
S3D Calc.
650 C betw een gridlines
700
Tw * [C] dT=10
Figure 6 Measured, and calculated temperature attenuation for Blades 2 and 3 and vanes 2 & 3.
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6. Measured metal temperature distribution of a cooled component, compared to
conjugate heat transfer calculations.
The metal temperatures of the components in the gas path are calculated using a 3D metal
model to which boundary conditions in terms of heat transfer coefficients and fluid
temperatures are coupled. Temperatures on the gas side are taken from the 3D NS calculations
described above, and the heat transfer coefficients are calculated using boundary layer
programmes. The cooling system inside the blade is described using a 1D flow network. Each
branch has correlations for heat transfer and pressure losses coupled to it. The calculated
temperatures are then transferred to the mechanical integrity model used for prediction of the
lifetime of the component. One of the main factors which determines the accuracy of the
lifetime prediction is how accurately the temperature distribution was predicted.
The main objective of the test was to measure the metal temperature distributions of the
components, in order to verify the temperature distributions from the design. If the
temperatures differ more than 20C between measurement and calculation, the calculation has
to be redone so that it simulates the measured temperatures, and in the worst case the cooling
has to be improved.
In this paper the thermal calculations for one of the cooled components, e.g. blade 2, is
presented in more detail. A photo of an instrumented blade 2 is shown in Fig 7. Each of the 3
identically instrumented blade 2 had 90 crystals.
The measured temperatures, the average of 3 blades, are shown in Figures 8, 9 and 10
together with results from the thermal calculations.
The scatter between different measured blades is satisfactorily low. The largest difference
found, (max- min) for a single point on the blade was 35 C and the average difference maxmin for similar points at different blades was 10 C.
Figure7. Blade 2 Instrumented with Thermo crystal
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800
81 820
0
770
820
780
760
750
750
74
0
0
74
730
790
750
740
730 720
760
740
720
0
730
77
710
750
760
720
730
770
770
740
790
780
710
750
770
760
780
760
760
710
720
770
730
770
750
740
720
730
770
710
750
730
770
720
740
760
780
715C
(-5)
694C
(+11)
753C
(+12)
750
730
71
720
0
740
74
0
760
750
76
0
750
751C
(-1)
0
76
761C
(-1)
710
715C
(+25)
20
74
0
740
717C
(-12)
700
750
760
780
770
750
750
740
740
760
750
760
780
750
760
740
740
744C
(-4)
0
730
0
73
740
720 710
700
730
71
7
750
71
0
770C
(0)
720
770
760
710
780
765C
(+20)
770
750
750
740
720
710
700
720
710
700
690
760
680
630
620
610
600
580
570
580
570
580
570
570
550
540
540
560
580
570
560
550
570
590
580
0
59
590
580
61
60 0
0
590
580
570
620
610
600
6200
61
600
580
740
730
Flow direction
723C
(-3)
742C
(+3)
770
760
750
740
730C
(+10)
710
780
753C
(+12)
743C
(-6)
742C
(+8)
720
752C
(+23)
770
770
760
750
770
760
750
759C
(-14)
722C
(-7)
737C
(+3)
770
760
762C
(+13)
770
710
746C
(+24)
0
78
789C
(-19)
760
730
800
775C
(+15)
750
785C
(-30)
767C
(-10)
7
760
786C
(-26)
767C
(+3)
725C
(-10)
10
720
75
0
739C
(+26)
80
0
790C
(-30)
73
0
720
560
530
530
520
Figure 8. Calculated and measured temperatures of blade 2 Suction side
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822C
(+3)
767C
(+23)
746C
(+34)
806C
(-1)
840
71
730
730
0
72
0
72 710
0
790
680
700
780
760
790
760
690
770
700
770
760
0
77
700
700
760
700
760
770
760
750
770
765C
(0)
75
0
720
710
700
760
750
750
752C
(+13)
760
760
770
760
710
750
760
740
0
69
743C
(+2)
760
750
770
0
74
750
750
740
730
700
760
770
776C
(-26)
760
690
760
0
750
730
765C
(-5)
700
720
710
730
749C
(-4)
74
0
753C
(-8)
0
76
761C
(-16)
0
73
748C
(-23)
690
0
72
720
760
750
740 0
73
720
710
788C
(-18)
770
770
785C
(-25)
75
740
748C
(-18)
720
710
710
770
760
690
760
770
0
750
74
770
760
690
710
780
760
800C
(-30)
740
730
700
720
710
780
770
760
770
750
740
720
710
700
690
740 750
730
760
780
760
750
740
730
730
757C
(-2)
770
796C
(-26)
800C
(-35)
774C
(-4)
730
720
710
700
690
770
744C
(0)
725C
(-14)
797C
(-17)
789C
(-19)
720
710
776C
(-6)
790
770
730
700
690
722C
(+3)
Flow direction
799C
(-39)
720
710
713C
(-3)
771C
(+4)
787C
(+3)
780
710
690
72
0
690
700
787C
(-12)
791C
(-31)
760
790
700
780
700
690
719C
(+6)
760
750
740
73
730
720
710
700
690
680
0
730
620
610
600
58
0
57
0
590
580
580
590
580
570
540
560
550
560
560
530
520
530
590
580
570
580
570
580
570
560
720
590
560
550
660
650
640
630
62
0
61
0
600
560
550
540
560
550
513C – Thermocrystal test
(-13) – Calc. result-measurement
Figure 9. Calculated and measured temperatures of blade 2 Pressure side
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721C
(-6)
5
72
720
76
0
755
705
0
76
5
75
0
75
5
74
0
74
5
73
68
0
758C
(+2)
0
760
77
68
0
750
785
73
0
740
72
5
710
705
715
830
730
0
72
740
735
740
0
750
720
74
5
74
0
73
5
73
0
72
5
75 755
0
780
820
730
72
0
745
740
735
730
725
750
715
71
0
74
0
73
5
750
745
740
735
730
755
760
5
73
744C
(+3)
729C
(+4)
715
825
725
710
740C
(+5)
745
740
735
740
810
805
720
5
735
5
75
74
740
5
750
5
71
770
73
5
74
5
0
74
73
0
72
5
72
0
731C
(+9)
75
0
745
74
5
67
690
765
760
751C
(+1)
775
765
0
69
700
695
0
73
0
77
685
79
690
765
770
765
657C
(+36)
77
5
733C
(+14)
513C – Thermocrystal test
(-13) – Calc. result-measurement
Figure 10 Calculated and measured temperatures of blade 2 platform
The results in Figs 8, 9 and 10 were obtained using gas side boundary conditions from the
local 3D aero model. In the local model the measured inlet gas temperature profile, Fig 6. has
been used. In Fig 8 which shows the suction side of the blade it is seen that there is an area at
mid-cord, that is calculated approximately 25 C too hot in the cooling model. On the pressure
side opposite to the area measured too cold the temperatures agree well between calculated
and measured. The blade is equipped with a multipass cooling system with ribbed channel at
leading edge and mid cord, and a race track type cooling system at the trailing edge. Close to
the tip at trailing edge the measured temperatures are higher than the calculated on both
pressure and suction side, by approximately 30-35 C. This indicates that the correlation for
the heat transfer coefficient is too optimistic on the inside of the blade. The temperature level
is, however, so low that this is not a problem. The hottest point measured on the blade is, not
very surprisingly, at the pressure side close to the tip 822C. The prediction of leading edge
temperatures is satisfactory; only two points differ more than 12C.
The prediction of the platform temperatures is very good; only one point differs more than
14C.That point is situated at the upstream part of the platform, close to the suction side. The
difference here comes from the difficulty in simulating how the purge flow from the cavity
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before the blade is distributed. The distribution is dependent on unsteady effects and local
geometry features that are not resolved in the degree of detail of the used3D aero model.
The measurement confirms the metal temperature level of blade 2 from the design project of
the engine.
7. Conclusion
After the successful performance of the test it must be stated that using thermo-crystals is a
reliable test method. It provides the possibility to measure temperatures in detail and pick up
the temperature gradients with good accuracy, particularly for rotating blades.
There is generally good correlation between measurements and calculations, which gives
confidence in the used calculation methods, and correlations. However, to ascertain a better
prediction of the gas temperature distributions in future the proposal is to use unsteady CFD
analysis as standard.
During the test several areas with potential for saving cooling air have been identified.
8. References
1. V.A Nikolaenko, V. I. Karnushin. Year 1986. Measurement of temperatures using
irradiated materials.
2. V.A Nikolaenko,V.A Morosov,N.I. Kasianov. Rev. int. Temp et Refract 1976 t.13 pp 17-20
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