01 GTP-13-1334

Gabriel Marinescue-mail: [email protected]

Wolfgang F. Mohre-mail: [email protected]

Andreas Ehrsame-mail: [email protected]

Paolo Ruffinoe-mail: [email protected]

Michael Selle-mail: [email protected]

Alstom, Power

Baden 5401, Switzerland

Experimental Investigation IntoThermal Behavior of SteamTurbine Components—Temperature MeasurementsWith Optical Probes and NaturalCooling AnalysisThe steam turbine cooldown has a significant impact on the cyclic fatigue life. A lower ini-tial metal temperature after standstill results in a higher temperature difference to be over-come during the next start-up. Generally, lower initial metal temperatures result in higherstart-up stress. In order to optimize steam turbines for cyclic operation, it is essential tofully understand natural cooling, which is especially challenging for rotors. This paperpresents a first-in-time application of a 2D numerical procedure for the assessment of thethermal regime during natural cooling, including the rotors, casings, valves, and mainpipes. The concept of the cooling calculation is to replace the fluid gross buoyancy duringnatural cooling by an equivalent fluid conductivity that gives the same thermal effect onthe metal parts. The fluid equivalent conductivity is calculated based on experimental data.The turbine temperature was measured with pyrometric probes on the rotor and withstandard thermocouples on inner and outer casings. The pyrometric probes were cali-brated with standard temperature measurements on a thermo well, where the steam trans-mittance and the rotor metal transmissivity were measured. [DOI: 10.1115/1.4025556]

Introduction

Modern steam turbines are operated at high pressure and tem-perature. In addition many steam power plants are today subjectto operation modes such as double shifts or load following opera-tion. Especially for combined cycle power plants and solar ther-mal plants fast start-up and high operational flexibility is required.

At base load operation the hot components are exposed tocreep. Additionally, high fatigue occurs because of the thermalstress during transient events such as start-up, shut down, or loadchanges. In order to design a fast starting and flexible steam tur-bine, the engineer deals with an important challenge due to thesensitivity of the cyclic lifetime assessment. The thermal stressarising in the hot thick-walled turbine components such as rotor,valves, and casings during turbine start-up is directly related tothe temperature gradient. The highest stress occurs when themachine ramps up from standstill to base load condition. For anaccurate thermal stress calculation the temperature profilebecomes a very important parameter. This paper presents amethod for the assessment of the thermal regime during naturalcooling of steam turbines.

Instrumentation With Optical Probes

An operational Alstom KA26-1 unit was instrumented withthree optical probes OT1, OT2, OT3; with 24 thermocouples typeN class A on inner casing; and with 40 thermocouples type Nclass A metal sheet protected on outer casing as presented onFig. 1. This was the first field turbine instrumentedwith optical pyrometers tracking the rotor temperature for almost

96 h. The inner casing during instrumentation at the AlstomMorelia—Mexico plant is presented on Fig. 2.

Alstom has developed in-house a flexible, fiber-based pyrome-ter [1–3] shown on Fig. 3. The flexible pyrometer consists of aprobe containing a low-OH gold-coated high temperature opticalglass fiber with a diameter of 0.3 mm and a numerical aperture of0.2. At the tip of the probe there is a sapphire lens of Ø2.4 mm,which reduces the numerical aperture of the system to 0.04. Thesignal picked up by the probe is then sent to an optical detector,an InGaAs PIN photodiode (three layer photo-diode with anintrinsec layer between the p- and n-type regions), G5853 ofHamamatsu. The photodiode is directly mounted to a compact pe-ripheral component interconnect card, which is based on a Motor-ola DSP56000 digital signal processor. The signal processor readsthe data of a 24 bit analog to digital converter with a sampling rateof 100,000 per second and converts the measured intensitydirectly into temperature. At temperatures above 230 �C, the tem-perature precision of the optical probe is better than 61.5 �C [1].Below this temperature, the precision quickly deteriorates and at150 �C reaches 610 �C. Below 130 �C the signal is useless as longas the irradiation signal vanishes in the dark current of thephotodiode.

Literature about the transmissivity of high-pressure stream isvery limited. Available papers and calculations are based on low-pressure data sets. This data highlights several transmittingwindows between strong absorption bands of steam, which aredetermined by the rotational and vibrational quantum states. Thelowest window W1 is located between 8 and 12 lm. At longerwavelengths the light is absorbed by pure rotational transitions,while at shorter wavelengths the light is absorbed by a rovibra-tional transition of the symmetric bend. The next windows rangefrom 3.5 to 4.3 lm (W2), from 2.0 to 2.4 lm (W3), and from 1.5to 1.7 lm (W4) (see Table 1).

Further, even more narrow consecutive windows exist towardshorter wavelengths. However, the blackbody radiation density

Contributed by the Controls, Diagnostics and Instrumentation Committee ofASME for publication in the JOURNAL OF ENGINEERING FOR GAS TURBINES AND POWER.Manuscript received August 31, 2013; final manuscript received September 10,2013; published online November 1, 2013. Editor: David Wisler.

Journal of Engineering for Gas Turbines and Power FEBRUARY 2014, Vol. 136 / 021602-1Copyright VC 2014 by ASME

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below 1.5 lm is too small to be used for high precision tempera-ture determination in the range from 250 �C up to 700 �C.

The steam transmittance is crucial for the intensity pyrometryin steam turbines. The usual approach [4] is to extrapolate thepressure broadening measured on low-pressure measurements asshown in Refs. [5–7]. This is dangerous, as low-pressure broaden-ing effects are dominated by two-body interactions, whereas high-pressure broadening effects are affected by many-body interac-tions or even spectral shifts caused by water clusters. Such many-body effects reduce the lifetime of the molecular rovibrationallevels, which further increases the pressure broadening. Also to beconsidered are line shape effects caused by the slow falloff of theLorenzian, Dicke [8], and Galatry [9] type. The slow falloff ofthese lines shapes leads at high pressure to a long-range artifact,where far from any line, like at 2.5 lm, the residual absorptionmay reach significant levels.

Measurements at high pressure are rare. We found absorptioncell measurements [10,11] and shock-tube results [12] discussingthe line shape but no measurements in the transmitting windows.Therefore, a dedicated autoclave was designed (see Fig. 4). Thesteam measured transmittance at 30 bar and 600 �C is shown onFig. 5 in comparison to extrapolated low resolution measurementsfrom Goldstein [10]. The program Spectralcalc was used to calcu-late the spectra. This program uses the line assignments of theHITRAN and HITEMP [13] database to calculate the strengths of

the lines as function of the temperature and uses the low-pressurebroadening data to linearly extrapolate the transmittance spectra tovery high pressures. The comparison between the theoretical pre-diction and the featured wavelengths shows good agreement. Theresults of the transmittance tests indicated that the intensity pyrome-try for IP and in particular for HP steam turbines is best conductedin the steam transmittance window W4 at wavelength 1.6 lm.

The optical probe lenses were protected against FeO particlescontamination with a nitrogen purge device. Figure 6 shows acomparison between a contaminated and purged lens in a realsteam turbine. This comparison confirmed that the purge wasmandatory to ensure measurement accuracy.

Measured Temperatures

The natural cooling measurements were conducted in Decem-ber 2010 during the power-plant commissioning phase. Themachine consists of a GT26 gas turbine and a HP-IP-LP steam tur-bine. Before starting the natural cooling measurements themachine was stabilized at base load regime. From base load thesteam turbine was by-passed and disconnected from the gasturbine. The glands system was maintained active together withvacuum in the turbine cavity for 3 h 15 min. After that the glandssystem was deactivated and ambient pressure established withinthe turbine cavity. The thermocouples and optical probe signalsrecorded the metal temperature for 96 h. After 96 h the machinewas ramped up to base load regime.

Figure 7 shows the transient temperatures measured by the opti-cal probes OT1, OT2, and OT3. The temperatures are given innondimensional format, divided by the live steam temperature atbase load. The temperatures where this ratio is below 0.35 reachedthe accuracy limit of the pyrometric method and weredisregarded.

Some of the temperatures recorded at the thermocouples loca-tion are presented on Fig. 8.

It must be noted that there are locations both on the inner andouter casing where the temperature increased within the first hoursafter natural cooling start. This phenomenon occurs on the colddomains once the active cooling specific for base load regime ends.

The Finite Element Analysis

The main difficulty of the natural cooling analysis consists ofthe long physical cooling time (approximately 100 h) relative tothe short integration time step (0.01 ms, typically) of the numericalscheme required for a convergent process. For this reason much

Fig. 2 The IP steam turbine arrangement

Fig. 1 IP steam turbine instrumentation

021602-2 / Vol. 136, FEBRUARY 2014 Transactions of the ASME


Fig. 4 The USC autoclave. On the left the full view, on the right the detail of the box.

Fig. 3 (a) Flexible pyrometric probe as used in gas turbine applications. (b) The measurementchain as used for the in-house developed pyrometer.

Table 1 Summary of the transmitting windows properties

Transmitting window W4 W3 W2 W1

Center wavelength (lm) 1.6 2.2 4.0 8.0Required dynamic range in bits 31 24 16 12Minimum temperature for equivalent noise temperature specified at 10 �C and for 1 Hz. 60 �C 40 �C <20 �C <20 �CMaximum operating temperature of optical fiber 700 700 130 70

Journal of Engineering for Gas Turbines and Power FEBRUARY 2014, Vol. 136 / 021602-3


attention was paid to the software used for modeling. Ideally thesoftware has to fulfill the following two conditions: (a) it has to beable to model the steam ingestion phase when the steam enthalpyfeeding the glands is distributed in the turbine cavity, and (b) it hasto be able to capture the thermal effect of the steam flow in the tur-bine cavity to transfer the heat from the hot rotor and inner casingto the outer casing, valves, pipes, and forward to ambient. One ofthe finite element applications qualified for these conditions isSC03, a Rolls-Royce in-house finite element software. AlstomPower and Rolls-Royce built a SC03 plug in for steam applicationsthat calculates automatically the steam thermodynamic propertiesand the corresponding heat transfer coefficients.

Consequently, a 2D transient SC03 model was built based onthe IP turbine geometry. The steam ingestion during the first 3 h15 min was modeled adding an assumed shape of the steam jetcontour. Figure 9 shows the jet contour and the location of thethermocouples T11.1, T24.1, and Tm33. The steam enthalpy wasgradually distributed from A to B along and inside the jet contour.

The numerical experiments showed that the position of thesteam jet contour in the turbine cavity has a negligible impact onthe metal temperature distribution. The most important is to bringthe steam glands energy in the turbine cavity distributed in time inline with the physical process, which is properly captured in the fi-nite element model. Condition (b) mentioned above was satisfiedintroducing finite elements in the turbine cavity defined with fluidconductivity (see Fig. 10).

The steam buoyancy, very active during the steam ingestionphase, can be interpreted as a heat wave that travels in the turbinecavity, driven by the local thermal gradient. The thermal effect ofthis buoyancy can be captured as a temperature-dependent con-ductivity, higher than a given reference fluid conductivity. Asmost of the time the natural cooling phase in the turbine cavity isair, we considered the air as the reference fluid. The thermal effectof the local buoyancy was captured via a correction factor K(T)introduced in the fluid conductivity k(T) [14]. Then, the fluid con-ductivity in the turbine cavity is

Fig. 6 Effect of purging on the lenses contamination in a real steam turbine (left not purged,right purged)

Fig. 5 The steam transmittance at 20 bar and 600 �C. The calculated curve using the HITRANdatabase [13], the low resolution data of Goldstein [10], and our experimental results from theFTIR spectrometer.



k Tð Þ¼ K Tð Þ�kair Tð Þ (1)

where kair(T) is the standard air conductivity and K(T)> 1. Physi-cally the function K(T) shows how many times the real heat wavein the turbine cavity travels faster than the air conductivity. K(T)was defined as a function of three parameters a1, a2, a3 used tomatch the thermal model relative to the experimental data.

KðTÞ ¼ a1T2 þ a2T þ a3 (2)

Consequently, the physical problem was reduced to the followingoptimization problem:

@ðT � TmeasÞ2

@a1

¼ 0;@ðT � TmeasÞ2

@a2

¼ 0;@ðT � TmeasÞ2

@a3

¼ 0 (3)

where T is the local metal temperature at coordinates x,y and thetime t with k(T) defined at (1). Tmeas is the time-dependent metaltemperature recorded on each thermocouple. The initial conditionwas set using a different model built with standard thermal bound-ary conditions that corresponds to the base load regime (seeFig. 11).

Equation (3) was solved iteratively starting from the followingremark. For each temperature T in the fluid domain there is acorrection factor K(p)(T) at iteration p and a K(pþ 1)(T) factor atiteration pþ 1 that underestimates, respectively, overestimates themeasured temperature Tmeas at each thermocouple location. T isthe metal temperature taken from the finite element modelcalculated at each thermocouple location. A linear interpolationbetween K(p)(T) and K(pþ 1)(T) gives the correction factor K(pþ 2)

(T) at the next iteration.

Kðpþ2Þ ¼ KðpÞ þ Tmeas�TðpÞ

Tðpþ1Þ � TðpÞ� Kðpþ1Þ � KðpÞ� �

(4)

Fig. 8 Inner and outer casing temperatures measured at T11.1,T24.1, Tm33, and Tm42

Fig. 10 Meshed model for natural cooling analysis

Fig. 11 Base load. Initial condition for natural cooling.

Fig. 7 Rotor temperature measured at optical probes OT1,OT2, and OT3

Fig. 9 Thermal boundary conditions to simulate the steamingestion



After each iteration the function K(T) was smoothed with the leastsquare method to compensate the scatter of the measured tempera-tures. The iterative process was ended once a norm of the vector (a1,a2, a3) became smaller than the method’s accuracy. The iterative pro-cess, presented in Fig. 12, was applied for each thermocouple gener-ating a corresponding Kj(T) function, where j is the thermocouple

index. At the end the Kj(T) functions were averaged (see Fig. 13).The averaged K(T) is called the “overconductivity function.”

It must be noted the large scatter of the Kj(T) functions (see Fig.13) at high temperatures (temperature/live steam temperature above0.50). This scatter, which defines the method accuracy, shows thatthe pressure gradient is not negligible at high temperatures.

Fig. 12 The iterative process

Fig. 13 The overconductivity function K(T)



Discussion on the 3D-2D Equivalence

As mentioned in the Finite Element Analysis section, only the2D models allow in a reasonable time the natural cooling analysis.The 2D models can simulate accurately the temperature on therotor and with an acceptable accuracy the temperature on the innerand outer casing. But the metal temperature distribution on thecasings, blades, valve, and feeding pipes are 3D, impacting thetemperature on 2D parts. That means an equivalence 3D-2D forthese parts is required in order to ensure the model accuracy.

The main idea of the 3D-2D equivalence is to redistribute uni-formly on circumference the mass of the 3D nonaxisymmetricdomains in such way to get in 2D a similar thermal effect. Oneach of these nonaxisymmetric domains a property called“thickness” was allocated. The thickness is calculated for eachspecific domain from the 3D to 2D mass equivalence condition.On each domain the corresponding thickness is constant but dif-ferent from domain to domain (see Fig. 14). The thickness is cal-culated for each specific domain from the mass equivalencecondition. Obviously this approach cannot have the accuracy of a3D analysis, but the numerical experiments showed that with acorrect selection of the thickness property, the model qualityremains acceptable. A special note for the inlet scroll—the cross-

section plan was selected in such a way that integrated on circum-ference, it gives the same mass as the 3D inner casing.

Results

The results show good agreement between the measured andcalculated temperatures. The model was calibrated within 15 �C…20 �C deviations versus measurements in order to remain conserv-ative. This conservatism is required to cover: (a) the uncertaintyof the steam ingestion mass flow and labyrinths deterioration, (b)the temperature deviations from machine to machine, and (c) the3D effects not captured in a 2D model (see Fig. 15). The thermalmodel captures properly the temperature increase on the cold partsduring the first hours of natural cooling, in this case on the outercasing (see Fig. 16). Figure 17 shows a comparison between themeasured and calculated metal temperature on the rotor at loca-tion OT1. Figures 18 and 19 show the calculated temperature mapat 2 h and 10 h, respectively, after the natural cooling start. Theresults suggest that the valve cools down slower than the rotorcore. This could be explained by the larger surface of the outercasing in contact with the ambient air and the contribution of thebuoyancy in the turbine cavity.

Once the FE model was calibrated, significant data can beextracted and interpreted, giving useful indication on the most im-portant parameters that impact the rotor cyclic life.

Figure 19 shows that at 10 h after natural cooling start, the ther-mal gradient within the valve is bigger than the thermal gradientwithin the turbine cavity. This seems to be the consequence of thesteam ingestion phase during the first 3 h 15 min after naturalcooling start. The overconductivity function K(T) captures prop-erly this effect increasing accordingly the fluid conductivity in theturbine cavity; meanwhile inside the valve the steam ingestion ismissing, so the temperature gradient drops down slower.

In order to analyze the impact of the above parameters wedefine the following time frames:

• Hot start (HS) condition corresponds to the turbine restart at8 h after natural cooling start.

• Warm start (WS) condition corresponds to the turbine restartat 60 h after natural cooling start.

Fig. 14 The “thickness” property

Fig. 15 Calculated and measured temperature at T11.1



If the machine cools down at different conditions relative tothe reference conditions, the temperature drops down differ-ently and the initial temperature for the next WS or HS is devi-ated in consequence. As example, let’s consider the referenceambient temperature 20 �C in the turbine enclosure. This condi-tion gives a reference temperature Tref(t) function of time (seeFig. 20). If the ambient temperature has a deviation from 20 �Cto 50 �C, the metal temperature at OT1 deviates from Tref(t) toT(t).

We introduce the temperature deviation DT as the differencebetween the T and Tref calculated at OT1. Obviously, DT is a func-tion of time and can be positive or negative.

DTðtÞ ¼ TðtÞ � TrefðtÞ (5)

Not only the ambient temperature impacts the deviation DT whenthe machine restarts. The impact of the following four parameterson DT was assessed:

Fig. 16 Calculated and measured temperature at Tm33

Fig. 17 Calculated and measured temperature at OT1



Fig. 18 Temperature distribution at 2 h after natural coolingstart

Fig. 19 Temperature distribution at 10 h after natural coolingstart

Fig. 20 Impact of the ambient temperature

Fig. 21 Impact of the different parameters on the rotor temperature at OT1



• bearing oil temperature¼ nominal tempþ (0 �C…40 �C)• ambient temperature¼ 20 �C…50 �C• ingested steam mass flow during the ingestion phase¼ 0%…

300% relative to the nominal ingested mass flow• HTC on the outer face of the outer casing¼ 70%…130%

relative to nominal HTC

Figure 21 collects the temperature deviation DT at OT1 foreach of the above parameters taken at 8 h (hot start condition),respectively, 60 h (warm start condition) after natural coolingstart.

Figure 21 suggests the following conclusions:

• The deviation of the bearing oil temperature relative to thestandard oil temperature has a low impact (1 deg…3 �C) onrotor temperature at WS and negligible at HS.

• The ambient temperature has a low impact (2 deg…3 �C) atHS and 16 deg…18 �C impact at WS.

• The deviation due to the steam ingestion has a 6 deg… 8 �Cimpact at HS and negligible at WS. The impact of the glandsteam temperature on rotor and casings was assessed inRef. [15] Sec. 5.3 and similar conclusions were found.

• The deviation of the thermal insulation quality has10 deg…12 �C impact at HS and a 30 �C…32 �C impact atthe WS.

Conclusions

A new numerical procedure for the assessment of the thermalregime during natural cooling of the main steam turbine compo-nents was validated with experimental measurements. Metal tem-peratures were measured on the rotor surface of a commercialsteam turbine with in-house developed pyrometers. Additionally alarge number of standard thermocouples were installed on theinner and outer casing.

The concept of the numerical cooling calculation is to replacethe fluid gross buoyancy during natural cooling by an equivalentfluid overconductivity that gives the same thermal effect on themetal parts. This fluid overconductivity function was establishedbased on experimental data.

The validation proved that the numerical model is able to pre-dict the cooling of all main steam turbine components with goodaccuracy. Based on the large number of metal temperature meas-urements available, the overall turbine cooling model was vali-dated. It was demonstrated that the numerical procedure is able tomodel the natural cooling heat transfer mechanism for 96 h physi-cal time on turbine rotor, casings, and valves. The calculationmethod, whose accuracy ranges within 0 deg…15 �C relative tomeasured data, was used to assess the impact of the physical pa-rameters the ambient air temperature, the steam ingestion time,the characteristics of thermal insulation, and the bearing oil tem-perature on the turbine rotor thermal regime.

The numerical cooling model can be used to provide importantinformation about the thermal state of the turbine parts during var-ious cooling events such as night shutdown, weekend shutdown,forced cooling events, etc. This is an important basis for the

design of flexible steam turbines, ready for fast and reliable cyclicoperation.

Nomenclature

a1, a2, a3 ¼ calibration parametersCCPP ¼ combined cycle power plant

HP ¼ high pressureHS ¼ hot start

HTC ¼ heat transfer coefficientIP ¼ intermediate pressure

K(T) ¼ correction factor for fluid conductivityp ¼ iteration numberT ¼ calculated metal temperature at a thermocouple

locationTfluid ¼ fluid temperatureTmeas ¼ measured metal temperature at a thermocouple

locationWS ¼ warm start

k ¼ fluid thermal conductivity

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