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Thermo-mechanical fatigue tests on uniaxial and component-like 1CrMoV rotor steel specimens
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International Journalof Fatigue
International Journal of Fatigue 30 (2008) 241–248
www.elsevier.com/locate/ijfatigue
Thermo-mechanical fatigue tests on uniaxial and component-like 1CrMoV rotor steel specimens F. Colombo a, E. Mazza a
a,b,*
, S.R. Holdsworth
b,c,1
, R.P. Skelton
d
Department of Mechanical Engineering, ETH Zurich, IMES, CLA J21.1, Tannenstrasse 3, CH-8092 Zurich, Switzerland b ¨ berlandstrasse 129, Dubendorf CH-8600, Switzerland EMPA - Materials Science & Technology, U c Alstom Power, Willans Works, Newbold Road, Rugby CV21 2NH, UK d Consultant, Guildford, UK Accepted 15 January 2007 Available online 19 March 2007
Abstract An extensive project has been undertaken to characterize the thermal fatigue behaviour of 1CrMoV rotor steel under service-like loading conditions. In the first phase of activity, thermo-mechanical fatigue (TMF) tests were performed on plain uniaxial tensile specimens subject to service-like loading conditions characterized by low strain rates, thermal transients to 565 C and long hold times at the peak temperature. These were followed by thermo-mechanical fatigue tests performed on circumferentially notched bars in order to study creep–fatigue crack initiation and short crack growth behaviour at stress concentrations. In the final phase of the project a new thermal fatigue experiment was implemented, applying severe thermal cycles to a large 3D feature specimen with a complex 10 mm deep groove. The experiment aimed to reproduce at specific notch locations loading conditions representative of turbine rotor features in service in terms of multiaxial stress state and thermal strain gradients. The results of these TMF and thermal fatigue studies are summarized in this paper. 2007 Elsevier Ltd. All rights reserved. Keywords: TMF; Thermal fatigue; Creep–fatigue interaction; Life prediction
1. Introduction The work presented in this paper is the result of a comprehensive research project aiming to improve the procedure for lifetime calculation of turbine components and to gather further understanding of the creep–fatigue damage interaction occurring during turbine operation. In particular, the creep–fatigue damage interaction at stress concentrations is analysed, for circumstances, where (i) the stress state is multiaxial, (ii) significant strain gradients are present, and (iii) the loading history is an-isothermal *
Corresponding author. Address: Department of Mechanical Engineering, ETH Zurich, IMES, CLA J21.1, Tannenstrasse 3, CH-8092 Zurich, Switzerland. Tel.: +41 (0)79 775 2381; fax: +41 (0)44 632 11 45. E-mail address: [email protected] (E. Mazza). 1 Formerly. 0142-1123/$ - see front matter 2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijfatigue.2007.01.036
and characterized by low strain rates (105/s) and significant hold periods at the peak cycle temperature. Traditionally, the risk of thermal fatigue cracking at such critical locations in components has been assessed using the results from isothermal tests conducted at (or close to) the peak operating temperature, see Timo [1], Mayer and Tremmel [2], Ha¨rkegard [3] and Dawson [4]. 1CrMoV steel is used widely for HP/IP steam turbine rotors operating at temperatures up to 565 C and there is extensive knowledge of its behaviour under isothermal LCF loading conditions without and with hold times, see Bhongbhibhat [5], Thomas and Dawson [6], Miller et al. [7] and Bicego et al. [8]. In the first phase of the project the behaviour of 1CrMoV steel subject to service-like thermo-mechanical fatigue (TMF) loading cycles was studied in uniaxial tests. Cyclic deformation behaviour and cycles to crack initiation
F. Colombo et al. / International Journal of Fatigue 30 (2008) 241–248
were determined with TMF cycles characterized by low strain rates, long hold periods at peak cycle temperature, and strain and temperature histories representative of the conditions experienced by component features during service, see Colombo et al. [9] and Masserey et al. [10]. The second phase of the research activity focussed on the creep–fatigue behaviour of circumferentially notched bars with isothermal fatigue (N-LCF) and thermo-mechanical fatigue (N-TMF) experiments, see Mazza et al. [11,12]. For the N-TMF tests the temperature and loading histories at the notch root were similar to the service-like TMF cycles applied in uniaxial tests. The cycles to crack initiation from these tests could therefore be directly compared with the results from the uniaxial tests. These results are reviewed in Section 3 of the paper. In the last phase of the project, a service-like thermal fatigue experiment with a component-like testpiece was investigated. This novel thermal fatigue test involved the application of cycles with severe thermal gradients to a large 3D feature specimen (approximately 200 · 200 · 200 mm) with a complex 10 mm deep groove. The experiment aimed to reproduce at specific notch locations loading conditions similar to those of features of a turbine rotor in service (in terms of multiaxial stress state and significant thermal strain gradients). Section 4 describes this experiment, the results in terms of crack development and the findings from post-test inspection. The conclusions are presented in Section 5. 2. Test material All testpieces were manufactured from the same heat of 1CrMoV rotor steel. The chemical composition and the Table 1 Material details (composition in weight %) C (%)
Cr (%)
Mo (%)
Ni (%)
V (%)
Rp0.2 (MPa)
RM (MPa)
0.25
0.88
0.76
0.69
0.33
640
780
600
mechanical properties of the material which originated from a high temperature HP steam turbine rotor forging are given in Table 1. The production forging had been oil quenched from 950/970 C and tempered at 695/700 C. The steel was characterized in terms of its tensile, creep and LCF (without and with hold time) properties, see Mazza et al. [11]. 3. Service-like TMF Tests 3.1. Uniaxial specimen tests The mechanical strain and temperature control profiles of the uniaxial TMF tests were selected to reproduce rotor service-like conditions. Temperature was varied between 350 and 565C with a heating rate of 4 or 14 C/min depending on the cycle type and a cooling rate of 18 C/ min (Fig. 1). The nominal mechanical strain range was 1.4%. Important features of the adopted TMF cycles were low strain rates (105/s) and long hold times at 565 C. During the first TMF tests, dimensional instability was observed in the form of gauge section necking. This was diagnosed and analytically demonstrated to be the result of a temperature variation along the testpiece gauge length which despite satisfying the requirements of existing TMF Standards was clearly unacceptable for this type of cycle, see Holdsworth et al. [13]. In subsequent tests, the testing procedure was changed to significantly reduce the magnitude of the temperature gradient along the testpiece gauge length and the consequent gauge section necking. Following a detailed study to fully characterize the thermal gradients existing during the first tests, see Holdsworth et al. [13], the necking phenomenon was investigated by finite element analysis with particular attention being paid to the local strain range amplification and its effect on cycles to crack initiation, see Masserey et al. [10]. A 2% load drop criterion was adopted as the crack initiation criterion for the data reported in Fig. 2.
0
600
0
temperature -3
550
-3
500
-6
500
-6
450
-9
strain 400
cycle with dwell
350
300 0
50
100
minutes
°C
550
strain (‰)
°C
temperature
-9
450
strain
-12
400
-15
350
-18
300
strain (‰)
242
-12
cycle w/o dwell
-15
-18 0
20
30
minutes
Fig. 1. Service-like TMF cycles. Left: cycle with 1 h hold time; right: continuous cycle.
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243
10
plain TMF w ith dw ell
strain range (%)
plain TMF w /o dw ell
N-LCF N-TMF 1 3D component-like LCF 565˚C 1 hour dw ell at
ε max
LCF 565˚C
location 1 location 2 0.1 10
100
1000
10000
N cycle
Fig. 2. Summary of endurance results: the reported strain ranges associated with locations 1 and 2 are effective values (see Section 4.5).
On the basis of isothermal tensile stress–strain and relaxation curves available for the specific heat of 1CrMoV steel under evaluation, a simple constitutive model (including temperature dependent equations for time-dependent and time-independent plasticity) was defined and demonstrated to be capable of effectively predicting the behaviour of the material in the uniaxial TMF tests, both with and without hold time. In both cases, an accurate matching between predicted and recorded force–displacement loops was achieved both for virgin and mid-life conditions (i.e. prior to the development of significant gauge section necking). The virgin and mid-life deformation behaviour of the material could be modelled with the variation of a single parameter, see Colombo et al. [9]. However, the model was also extended to describe the evolutionary behaviour from virgin to mid-life, as described in Mazza et al. [12]. Post-test inspection was carried out on all testpieces to reveal the consecutive development of two damage mechanisms: (i) fatigue at the surface and (ii) creep in the interior, see Holdsworth et al. [13]. 3.2. Notched specimen tests LCF and TMF tests were performed with circumferentially notched bars. The remote displacement and temperature control profiles of the N-TMF tests were defined on the basis of preliminary finite element (FE) calculations in order to produce similar loading histories at the notch root to the uniaxial TMF tests described in Section 3.1. The main purpose of the investigation was to evaluate the suitability of using uniaxial data for the design life
assessment of components containing stress concentrations. In Fig. 2 the N-LCF and N-TMF test cycles to crack initiation are reported and can be compared with the isothermal 565 C LCF data lines for this steel, see Mazza et al. [12]. The notched specimen test cycles to crack initiation are represented by a pair of triangles plotted at the corresponding notch root strain ranges. For each pair of triangles, the filled points represent the cycles to a crack size of 0.1 mm and the open points represent the cycles to a crack size of 0.5 mm. Life expressed as cycles to crack initiation requires the definition of an initiation criterion (corresponding to a critical size/condition at which the presence of the crack is first acknowledged). This is because the number of cycles to crack initiation is sensitive to the adopted initiation criterion in particular for notched specimen tests in which crack growth rates progressively decrease due to the reducing strain gradient ahead of the notch root. Good agreement between uniaxial and notched specimen cycles to crack initiation was obtained for a crack initiation criterion of 0.1 mm. Post-test metallographic inspection of the N-LCF and N-TMF specimens indicated that cracks were fully circumferential by the time they had penetrated to a depth of 0.7 mm. Crack development from the notch root in the N-LCF test was fully transgranular with no evidence of creep damage. Crack development in the N-TMF tests was also predominantly fatigue but with some evidence of associated creep damage. The development of damage in the notched specimens was notably different to that in the uniaxial specimens. While the applied mechanical strain range at the groove root of the notched specimens was
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identical to that in the gauge section of the uniaxial testpieces, the reference stress in the net section of the notched specimens was significantly lower than the net section stress in the uniaxial testpieces. There was therefore not the same driving force for creep damage development in the net section of the notched testpieces.
reproduce at specific notch locations loading conditions representative of turbine rotor features in service in terms of (i) multiaxial stress state, (ii) thermal strain gradient and (iii) an an-isothermal loading history involving low strain rates (105–107/s) and long hold times at high temperature (560 C).
4. 3D specimen tests
4.1. Testpiece
In the final phase of the investigation, thermal fatigue tests were performed on a large 1CrMoV 3D feature specimen with a complex 10 mm deep groove. The test aimed to
The large 3D feature specimen was manufactured from the same heat of 1CrMoV steel used for the uniaxial plain and notched specimen tests. The overall shape of the test-
Fig. 3. 3D feature specimen. (a) front view; (b) view of the bottom notched face; (c) view of the upper surface.
F. Colombo et al. / International Journal of Fatigue 30 (2008) 241–248
piece and the local notch geometry (Fig. 3) were optimized through iterative FE calculations, in order to achieve at specific notch locations severe temperature and strain transients. A complex 10 mm deep notch with two different tip radii of 3 and 5 mm was machined into the bottom face of the testpiece, Fig. 3. By inserting two through-section slots orthogonally to the main imposed heat flux direction, a strong enhancement of the generated temperature gradients could be achieved. 4.2. Experimental details For these experiments a special furnace was developed in which thermal cycles involving high temperature transients were applied to the bottom face of the testpiece by means of a base plate with embedded electrical heating elements and water cooled channels. The system enabled the specified heating and cooling rates to be achieved during the start-up, hold time and shut-down phases of the thermal cycle. The imposed thermal cycles at the bottom surface comprised: (i) 20 min of fast heating at a rate of 15–18 C/min followed by (ii) a dwell time at Tmax (560 C) with a duration of 3.5 h and (iii) forced cooling at a rate of 5–8 C/min to close the cycle, see Fig. 4a. The temperature gradients induced in the vertical direction (away from bottom surface of the testpiece) were recorded in each cycle by means of temperature measurements at a number of locations. To this end, 21 sheathed type K thermocouples were embedded in
245
the testpiece at different positions and connected to a data acquisition system. 4.3. Finite element calculations The measurements acquired by the thermocouples provided the basis for the FE representation of the transient temperature distribution within the testpiece during thermal cycling. Through suitable selection of the thermal boundary conditions (in terms of temperatures and heat transfer coefficients) an accurate matching between measured and predicted temperature histories within the testpiece was achieved. An illustration of the predicted temperature field at the end of the heating ramp by the thermal simulation is shown in Fig. 5a. The results of the temperature calculation were used as the input for the transient stress analysis. For this calculation, the heat-specific constitutive model validated with the experiments described in Section 3 was implemented. An example of the calculated distribution of equivalent (Mises) stress at the end of the heating ramp is shown in Fig. 5b. Larger strain ranges were determined at the 3 mm radius notch tip (location 1) and at the 5 mm radius notch tip (location 2), as indicated in Fig. 5b. Examples of predicted equivalent stress–strain loop and local strain and temperature histories at location 1 are reported in Fig. 4b and c for cycle number 40. Noteworthy are the high levels of hydrostatic stress (stress triaxiality) and the constrained stress relaxation during the hold times.
temperature ( °C)
600 500 400 300 200 100
0
100
200
300
400
500
time (min) 600 hydrostatic stress T = 540 ˚C
-3.0
-2.0
-1.0
0.0
400
0 -1
0
100
200
300
400
-3
-600 T = 320 ˚C
equivalent stress
total strain%
plastic strain%
100 0
-4
mech strain%
500300
200
-2
-400
-800
1.0
600 500
creep strain%
1
0 -200
temperature
2
strain %
stress (MPa)
400 200
3
T = 390 ˚C
T = 560 ˚C
time (min)
Fig. 4. 3D component-like experiment, cycle 40: (a) temperature history at the bottom face; (b) FE predicted stress–strain loops and (c) strain and temperature histories at location 1.
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The thickness of the oxide scale on the thermal fatigue crack faces seemed high for the time spent at high temperature (Fig. 6b). The applied thermal transients had clearly been responsible for enhanced spallation activity and this will have influenced oxidation kinetics and possibly cycles to crack initiation.
a
4.5. Thermal history
b
c total strain range (%)
4.0
N = 50 multiple cracks detected with MPI at location 1
3.0
2.0 1.0
N = 70 multiple cracks detected with MPI at location 2
location 1 location 2
0.0
0
10
20
30 40 50 cycle number
60
70
80
Fig. 5. FE calculations: (a) temperature and (b) equivalent stress distribution at the end of the heating ramp. Critical locations 1 and 2 are indicated. (c) Total strain range history along the specimen life.
4.4. Inspection results The 3D feature specimen was examined every 10 cycles by magnetic particle (MPI) and replica inspection. Multiple cracking was observed at location 1 after 50 cycles, with a crack network having developed in the 3 mm radius notch, growing in a direction parallel to the groove axis. The total width of the crack network after 50 cycles was 20 mm. Cracking was first observed at location 2 after 70 cycles (see Fig. 6a) and well developed cracks were observed at the end of the test, i.e. after 115 cycles. Images from the post-test inspection are shown in Fig. 6b, crack development at locations 1 and 2 being driven predominantly by fatigue but with evidence of consequential creep damage associated with the crack path (Fig. 6b).
The generation of identical thermal fatigue cycles for the entire test duration proved not to be possible for this experimental arrangement. From time to time, changes to the thermal boundary conditions occurred during the course of the test. The changes were primarily due to (i) periodical testpiece inspections during the course of the experiment, requiring the removal and repositioning of the testpiece on the heating/cooling elements, causing changes in the contact and heat transfer conditions, and (ii) thermal fatigue cracking of the heating/cooling elements requiring weld repair leading to geometry modifications to the elements. FE analysis of the temperature field history matching the measurements of each thermal fatigue cycle block were performed and used as input for the calculation of the stress and strain histories. Thus, the evolution of local notch root strain ranges during the course of the test history was quantified. For the reasons given above, the notch root strain ranges varied in a non systematic way between 2.1 and 3.2% for location 1 and between 1.4 and 2.4% for location 2, Fig. 5c. In order to rationalize this variable loading history, not untypical of that encountered by real turbine components in service, effective strain ranges for locations 1 and 2 were calculated for each loading block throughout specimen life until the visual appearance of cracking, i.e. using X Deeff ¼ ðN j Dej Þ=N i ð1Þ j
where Deeff is the effective strain range, Nj is the number of cycles spent at strain range Dej in the jth loading block and Ni is the number of cycles to the observation of visual cracking (see Section 4.4). The effective strain ranges determined according to (1) were 2.65% for location 1 at 50 cycles and 1.8% for location 2 at 70 cycles. The 3D specimen Ni(Deeff) results for the two notch root locations are significantly shorter than a life prediction based on the 565 C LCF with 1 h hold time data line (Fig. 2). This is primarily due to the higher creep damage generated at the notch radii as a consequence of the constrained stress relaxation during the hold time at the peak temperature in the thermal cycle (Fig. 4b). The results of all tests are rationalized by a formal creep–fatigue damage assessment analysis. The sensitivity of the creep–fatigue damage summation locus for 1CrMoV rotor steel to different deformation and damage calculation solutions applied to the TMF and thermal fatigue test results is under evaluation in a similar way to that performed in the BE 5245 project [14].
F. Colombo et al. / International Journal of Fatigue 30 (2008) 241–248
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Fig. 6. Inspection at the 3D feature specimen: (a) Image from MPI inspection at cycle N = 70 with crack network at locations 1 and 2. (b) Typical crack appearance at locations 1 and 2 at low (left) and high (right) magnifications: pictures from the post-test inspection (N = 115 cycles).
5. Conclusions
References
The thermal fatigue behaviour of uniaxial and a range of component-like 1CrMoV rotor steel specimens under service-like loading conditions has been investigated. The results provide valuable information on the cyclic deformation behaviour and cycles to crack initiation characteristics of 1CrMoV rotor steel subject to service-like conditions. In addition, they serve to validate the effectiveness of a previously established constitutive model based on isothermal material property data. In addition, the following conclusions are also drawn. The cycles to crack initiation in isothermal and servicecycle thermo-mechanical fatigue tests on uniaxial and circumferentially notched round bar specimens can be rationalized when the cycles involve hold times of similar duration at the same temperature and a consistent crack initiation criterion is adopted (in this case, a discernable crack extension increment of 0.1 mm). The simple use of isothermal LCF with 1 h hold time endurance data cannot be used in isolation to predict cycles to thermal fatigue crack initiation, in particular in circumstances where stress relaxation is constrained by high triaxiality in long hold times (e.g. as in the 3D specimen thermal fatigue tests). In such cases, the prediction of component life requires a more comprehensive assessment of creep and fatigue damage.
[1] Timo DP. Designing turbine components for low-cycle fatigue. In: Proceedings of the international conference on thermal stresses and fatigue, Berkeley; 1969. p. 453–69. [2] Mayer K-H, Tremmel D. The thermal fatigue of components in steam power plant. In: Proceedings of the international dvm symposium on low cycle fatigue strength and elasto-plastic behaviour of materials, Stuttgart, 8/9 October 1979. p. 105–16. [3] Ha¨rkegard G. Designing steam turbines for transient loading. In: Larsson LH, editor. High temperature structural design, ESIS12. Mechanical Engineering Publications; 1992. p. 21–40. [4] Dawson RAT. Monitoring and control of thermal stress and component life expenditure in steam turbines. In: Proceedings of the international conference on modern power stations, AIM, Lie`ge, September 1989. [5] Bhongbhibhat S. Untersuchungen u¨ber das Werkstoffverhalten in Gebiet der Zeitfestigkeit zur Erstellung von Berechnungsunterlagen fu¨r Uberwiegend Thermisch Beanspruchte Bauteille. Technischerwissenshaflicher Bericht, MPA Stuttgart, Heft 79-02, 1979. [6] Thomas G, Dawson RAT. The effect of dwell period and cycle type on high strain fatigue properties of 1CrMoV rotor forgings at 500–550 C. In: Proceedings of the international mechanical engineering conference on engineering aspects of creep, Sheffield, 1980. Paper C335/80. [7] Miller DA, Priest RH, Ellison EG. A review of material response and life prediction techniques under fatigue–creep loading conditions. High Temp Mater Process 1984;6(3, 4):155–94. [8] Bicego V, Fossati C, Ragazzoni S. Low cycle fatigue characterisation of a HP-IP steam turbine rotor. Low cycle fatigue. ASTM STP, vol. 942. West Conshohocken, PA: ASTM International; 1988. p. 1237–60. [9] Colombo F, Masserey B, Mazza E, Holdsworth SR. Simple modelling of the constitutive behaviour of a 1%CrMoV rotor steel in service-like thermo-mechanical fatigue tests. Mater High Temp 2002;19(4):223–5. [10] Masserey B, Colombo F, Mazza E, Holdsworth SR. Factors influencing the service-like thermomechanical fatigue test cycle
Acknowledgment The financial support of Alstom Power, Steam Turbines R&D, is gratefully acknowledged.
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endurance of 1%CrMoV rotor steel. Fatigue Fract Eng Mater Struct 2003;26(11):1041–52. [11] Mazza E, Holdsworth SR, Skelton RP. Characterization of the creep–fatigue behaviour of a 1CrMoV turbine steel. Mater High Temp 2004;21(3):119–28. [12] Mazza E, Hollenstein M, Holdsworth SR, Skelton RP. Notched specimens thermo-mechanical fatigue of a 1CrMoV turbine steel. Nucl Eng Design 2004;234:11–24.
[13] Holdsworth SR, Mazza E, Jung A. The response of 1CrMoV rotor steel to service-cycle thermo-mechanical fatigue testing. J Test Eval 2004;32(4):255–61. [14] Brite-Euram project BE 5245. Optimisation of methodologies to predict crack initiation and early growth in components under complex creep–fatigue loading (C-FAT). In: Holdsworth, editor. Synthesis report for publication. 1997.
International Journalof Fatigue
International Journal of Fatigue 30 (2008) 241–248
www.elsevier.com/locate/ijfatigue
Thermo-mechanical fatigue tests on uniaxial and component-like 1CrMoV rotor steel specimens F. Colombo a, E. Mazza a
a,b,*
, S.R. Holdsworth
b,c,1
, R.P. Skelton
d
Department of Mechanical Engineering, ETH Zurich, IMES, CLA J21.1, Tannenstrasse 3, CH-8092 Zurich, Switzerland b ¨ berlandstrasse 129, Dubendorf CH-8600, Switzerland EMPA - Materials Science & Technology, U c Alstom Power, Willans Works, Newbold Road, Rugby CV21 2NH, UK d Consultant, Guildford, UK Accepted 15 January 2007 Available online 19 March 2007
Abstract An extensive project has been undertaken to characterize the thermal fatigue behaviour of 1CrMoV rotor steel under service-like loading conditions. In the first phase of activity, thermo-mechanical fatigue (TMF) tests were performed on plain uniaxial tensile specimens subject to service-like loading conditions characterized by low strain rates, thermal transients to 565 C and long hold times at the peak temperature. These were followed by thermo-mechanical fatigue tests performed on circumferentially notched bars in order to study creep–fatigue crack initiation and short crack growth behaviour at stress concentrations. In the final phase of the project a new thermal fatigue experiment was implemented, applying severe thermal cycles to a large 3D feature specimen with a complex 10 mm deep groove. The experiment aimed to reproduce at specific notch locations loading conditions representative of turbine rotor features in service in terms of multiaxial stress state and thermal strain gradients. The results of these TMF and thermal fatigue studies are summarized in this paper. 2007 Elsevier Ltd. All rights reserved. Keywords: TMF; Thermal fatigue; Creep–fatigue interaction; Life prediction
1. Introduction The work presented in this paper is the result of a comprehensive research project aiming to improve the procedure for lifetime calculation of turbine components and to gather further understanding of the creep–fatigue damage interaction occurring during turbine operation. In particular, the creep–fatigue damage interaction at stress concentrations is analysed, for circumstances, where (i) the stress state is multiaxial, (ii) significant strain gradients are present, and (iii) the loading history is an-isothermal *
Corresponding author. Address: Department of Mechanical Engineering, ETH Zurich, IMES, CLA J21.1, Tannenstrasse 3, CH-8092 Zurich, Switzerland. Tel.: +41 (0)79 775 2381; fax: +41 (0)44 632 11 45. E-mail address: [email protected] (E. Mazza). 1 Formerly. 0142-1123/$ - see front matter 2007 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijfatigue.2007.01.036
and characterized by low strain rates (105/s) and significant hold periods at the peak cycle temperature. Traditionally, the risk of thermal fatigue cracking at such critical locations in components has been assessed using the results from isothermal tests conducted at (or close to) the peak operating temperature, see Timo [1], Mayer and Tremmel [2], Ha¨rkegard [3] and Dawson [4]. 1CrMoV steel is used widely for HP/IP steam turbine rotors operating at temperatures up to 565 C and there is extensive knowledge of its behaviour under isothermal LCF loading conditions without and with hold times, see Bhongbhibhat [5], Thomas and Dawson [6], Miller et al. [7] and Bicego et al. [8]. In the first phase of the project the behaviour of 1CrMoV steel subject to service-like thermo-mechanical fatigue (TMF) loading cycles was studied in uniaxial tests. Cyclic deformation behaviour and cycles to crack initiation
F. Colombo et al. / International Journal of Fatigue 30 (2008) 241–248
were determined with TMF cycles characterized by low strain rates, long hold periods at peak cycle temperature, and strain and temperature histories representative of the conditions experienced by component features during service, see Colombo et al. [9] and Masserey et al. [10]. The second phase of the research activity focussed on the creep–fatigue behaviour of circumferentially notched bars with isothermal fatigue (N-LCF) and thermo-mechanical fatigue (N-TMF) experiments, see Mazza et al. [11,12]. For the N-TMF tests the temperature and loading histories at the notch root were similar to the service-like TMF cycles applied in uniaxial tests. The cycles to crack initiation from these tests could therefore be directly compared with the results from the uniaxial tests. These results are reviewed in Section 3 of the paper. In the last phase of the project, a service-like thermal fatigue experiment with a component-like testpiece was investigated. This novel thermal fatigue test involved the application of cycles with severe thermal gradients to a large 3D feature specimen (approximately 200 · 200 · 200 mm) with a complex 10 mm deep groove. The experiment aimed to reproduce at specific notch locations loading conditions similar to those of features of a turbine rotor in service (in terms of multiaxial stress state and significant thermal strain gradients). Section 4 describes this experiment, the results in terms of crack development and the findings from post-test inspection. The conclusions are presented in Section 5. 2. Test material All testpieces were manufactured from the same heat of 1CrMoV rotor steel. The chemical composition and the Table 1 Material details (composition in weight %) C (%)
Cr (%)
Mo (%)
Ni (%)
V (%)
Rp0.2 (MPa)
RM (MPa)
0.25
0.88
0.76
0.69
0.33
640
780
600
mechanical properties of the material which originated from a high temperature HP steam turbine rotor forging are given in Table 1. The production forging had been oil quenched from 950/970 C and tempered at 695/700 C. The steel was characterized in terms of its tensile, creep and LCF (without and with hold time) properties, see Mazza et al. [11]. 3. Service-like TMF Tests 3.1. Uniaxial specimen tests The mechanical strain and temperature control profiles of the uniaxial TMF tests were selected to reproduce rotor service-like conditions. Temperature was varied between 350 and 565C with a heating rate of 4 or 14 C/min depending on the cycle type and a cooling rate of 18 C/ min (Fig. 1). The nominal mechanical strain range was 1.4%. Important features of the adopted TMF cycles were low strain rates (105/s) and long hold times at 565 C. During the first TMF tests, dimensional instability was observed in the form of gauge section necking. This was diagnosed and analytically demonstrated to be the result of a temperature variation along the testpiece gauge length which despite satisfying the requirements of existing TMF Standards was clearly unacceptable for this type of cycle, see Holdsworth et al. [13]. In subsequent tests, the testing procedure was changed to significantly reduce the magnitude of the temperature gradient along the testpiece gauge length and the consequent gauge section necking. Following a detailed study to fully characterize the thermal gradients existing during the first tests, see Holdsworth et al. [13], the necking phenomenon was investigated by finite element analysis with particular attention being paid to the local strain range amplification and its effect on cycles to crack initiation, see Masserey et al. [10]. A 2% load drop criterion was adopted as the crack initiation criterion for the data reported in Fig. 2.
0
600
0
temperature -3
550
-3
500
-6
500
-6
450
-9
strain 400
cycle with dwell
350
300 0
50
100
minutes
°C
550
strain (‰)
°C
temperature
-9
450
strain
-12
400
-15
350
-18
300
strain (‰)
242
-12
cycle w/o dwell
-15
-18 0
20
30
minutes
Fig. 1. Service-like TMF cycles. Left: cycle with 1 h hold time; right: continuous cycle.
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243
10
plain TMF w ith dw ell
strain range (%)
plain TMF w /o dw ell
N-LCF N-TMF 1 3D component-like LCF 565˚C 1 hour dw ell at
ε max
LCF 565˚C
location 1 location 2 0.1 10
100
1000
10000
N cycle
Fig. 2. Summary of endurance results: the reported strain ranges associated with locations 1 and 2 are effective values (see Section 4.5).
On the basis of isothermal tensile stress–strain and relaxation curves available for the specific heat of 1CrMoV steel under evaluation, a simple constitutive model (including temperature dependent equations for time-dependent and time-independent plasticity) was defined and demonstrated to be capable of effectively predicting the behaviour of the material in the uniaxial TMF tests, both with and without hold time. In both cases, an accurate matching between predicted and recorded force–displacement loops was achieved both for virgin and mid-life conditions (i.e. prior to the development of significant gauge section necking). The virgin and mid-life deformation behaviour of the material could be modelled with the variation of a single parameter, see Colombo et al. [9]. However, the model was also extended to describe the evolutionary behaviour from virgin to mid-life, as described in Mazza et al. [12]. Post-test inspection was carried out on all testpieces to reveal the consecutive development of two damage mechanisms: (i) fatigue at the surface and (ii) creep in the interior, see Holdsworth et al. [13]. 3.2. Notched specimen tests LCF and TMF tests were performed with circumferentially notched bars. The remote displacement and temperature control profiles of the N-TMF tests were defined on the basis of preliminary finite element (FE) calculations in order to produce similar loading histories at the notch root to the uniaxial TMF tests described in Section 3.1. The main purpose of the investigation was to evaluate the suitability of using uniaxial data for the design life
assessment of components containing stress concentrations. In Fig. 2 the N-LCF and N-TMF test cycles to crack initiation are reported and can be compared with the isothermal 565 C LCF data lines for this steel, see Mazza et al. [12]. The notched specimen test cycles to crack initiation are represented by a pair of triangles plotted at the corresponding notch root strain ranges. For each pair of triangles, the filled points represent the cycles to a crack size of 0.1 mm and the open points represent the cycles to a crack size of 0.5 mm. Life expressed as cycles to crack initiation requires the definition of an initiation criterion (corresponding to a critical size/condition at which the presence of the crack is first acknowledged). This is because the number of cycles to crack initiation is sensitive to the adopted initiation criterion in particular for notched specimen tests in which crack growth rates progressively decrease due to the reducing strain gradient ahead of the notch root. Good agreement between uniaxial and notched specimen cycles to crack initiation was obtained for a crack initiation criterion of 0.1 mm. Post-test metallographic inspection of the N-LCF and N-TMF specimens indicated that cracks were fully circumferential by the time they had penetrated to a depth of 0.7 mm. Crack development from the notch root in the N-LCF test was fully transgranular with no evidence of creep damage. Crack development in the N-TMF tests was also predominantly fatigue but with some evidence of associated creep damage. The development of damage in the notched specimens was notably different to that in the uniaxial specimens. While the applied mechanical strain range at the groove root of the notched specimens was
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identical to that in the gauge section of the uniaxial testpieces, the reference stress in the net section of the notched specimens was significantly lower than the net section stress in the uniaxial testpieces. There was therefore not the same driving force for creep damage development in the net section of the notched testpieces.
reproduce at specific notch locations loading conditions representative of turbine rotor features in service in terms of (i) multiaxial stress state, (ii) thermal strain gradient and (iii) an an-isothermal loading history involving low strain rates (105–107/s) and long hold times at high temperature (560 C).
4. 3D specimen tests
4.1. Testpiece
In the final phase of the investigation, thermal fatigue tests were performed on a large 1CrMoV 3D feature specimen with a complex 10 mm deep groove. The test aimed to
The large 3D feature specimen was manufactured from the same heat of 1CrMoV steel used for the uniaxial plain and notched specimen tests. The overall shape of the test-
Fig. 3. 3D feature specimen. (a) front view; (b) view of the bottom notched face; (c) view of the upper surface.
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piece and the local notch geometry (Fig. 3) were optimized through iterative FE calculations, in order to achieve at specific notch locations severe temperature and strain transients. A complex 10 mm deep notch with two different tip radii of 3 and 5 mm was machined into the bottom face of the testpiece, Fig. 3. By inserting two through-section slots orthogonally to the main imposed heat flux direction, a strong enhancement of the generated temperature gradients could be achieved. 4.2. Experimental details For these experiments a special furnace was developed in which thermal cycles involving high temperature transients were applied to the bottom face of the testpiece by means of a base plate with embedded electrical heating elements and water cooled channels. The system enabled the specified heating and cooling rates to be achieved during the start-up, hold time and shut-down phases of the thermal cycle. The imposed thermal cycles at the bottom surface comprised: (i) 20 min of fast heating at a rate of 15–18 C/min followed by (ii) a dwell time at Tmax (560 C) with a duration of 3.5 h and (iii) forced cooling at a rate of 5–8 C/min to close the cycle, see Fig. 4a. The temperature gradients induced in the vertical direction (away from bottom surface of the testpiece) were recorded in each cycle by means of temperature measurements at a number of locations. To this end, 21 sheathed type K thermocouples were embedded in
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the testpiece at different positions and connected to a data acquisition system. 4.3. Finite element calculations The measurements acquired by the thermocouples provided the basis for the FE representation of the transient temperature distribution within the testpiece during thermal cycling. Through suitable selection of the thermal boundary conditions (in terms of temperatures and heat transfer coefficients) an accurate matching between measured and predicted temperature histories within the testpiece was achieved. An illustration of the predicted temperature field at the end of the heating ramp by the thermal simulation is shown in Fig. 5a. The results of the temperature calculation were used as the input for the transient stress analysis. For this calculation, the heat-specific constitutive model validated with the experiments described in Section 3 was implemented. An example of the calculated distribution of equivalent (Mises) stress at the end of the heating ramp is shown in Fig. 5b. Larger strain ranges were determined at the 3 mm radius notch tip (location 1) and at the 5 mm radius notch tip (location 2), as indicated in Fig. 5b. Examples of predicted equivalent stress–strain loop and local strain and temperature histories at location 1 are reported in Fig. 4b and c for cycle number 40. Noteworthy are the high levels of hydrostatic stress (stress triaxiality) and the constrained stress relaxation during the hold times.
temperature ( °C)
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Fig. 4. 3D component-like experiment, cycle 40: (a) temperature history at the bottom face; (b) FE predicted stress–strain loops and (c) strain and temperature histories at location 1.
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The thickness of the oxide scale on the thermal fatigue crack faces seemed high for the time spent at high temperature (Fig. 6b). The applied thermal transients had clearly been responsible for enhanced spallation activity and this will have influenced oxidation kinetics and possibly cycles to crack initiation.
a
4.5. Thermal history
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Fig. 5. FE calculations: (a) temperature and (b) equivalent stress distribution at the end of the heating ramp. Critical locations 1 and 2 are indicated. (c) Total strain range history along the specimen life.
4.4. Inspection results The 3D feature specimen was examined every 10 cycles by magnetic particle (MPI) and replica inspection. Multiple cracking was observed at location 1 after 50 cycles, with a crack network having developed in the 3 mm radius notch, growing in a direction parallel to the groove axis. The total width of the crack network after 50 cycles was 20 mm. Cracking was first observed at location 2 after 70 cycles (see Fig. 6a) and well developed cracks were observed at the end of the test, i.e. after 115 cycles. Images from the post-test inspection are shown in Fig. 6b, crack development at locations 1 and 2 being driven predominantly by fatigue but with evidence of consequential creep damage associated with the crack path (Fig. 6b).
The generation of identical thermal fatigue cycles for the entire test duration proved not to be possible for this experimental arrangement. From time to time, changes to the thermal boundary conditions occurred during the course of the test. The changes were primarily due to (i) periodical testpiece inspections during the course of the experiment, requiring the removal and repositioning of the testpiece on the heating/cooling elements, causing changes in the contact and heat transfer conditions, and (ii) thermal fatigue cracking of the heating/cooling elements requiring weld repair leading to geometry modifications to the elements. FE analysis of the temperature field history matching the measurements of each thermal fatigue cycle block were performed and used as input for the calculation of the stress and strain histories. Thus, the evolution of local notch root strain ranges during the course of the test history was quantified. For the reasons given above, the notch root strain ranges varied in a non systematic way between 2.1 and 3.2% for location 1 and between 1.4 and 2.4% for location 2, Fig. 5c. In order to rationalize this variable loading history, not untypical of that encountered by real turbine components in service, effective strain ranges for locations 1 and 2 were calculated for each loading block throughout specimen life until the visual appearance of cracking, i.e. using X Deeff ¼ ðN j Dej Þ=N i ð1Þ j
where Deeff is the effective strain range, Nj is the number of cycles spent at strain range Dej in the jth loading block and Ni is the number of cycles to the observation of visual cracking (see Section 4.4). The effective strain ranges determined according to (1) were 2.65% for location 1 at 50 cycles and 1.8% for location 2 at 70 cycles. The 3D specimen Ni(Deeff) results for the two notch root locations are significantly shorter than a life prediction based on the 565 C LCF with 1 h hold time data line (Fig. 2). This is primarily due to the higher creep damage generated at the notch radii as a consequence of the constrained stress relaxation during the hold time at the peak temperature in the thermal cycle (Fig. 4b). The results of all tests are rationalized by a formal creep–fatigue damage assessment analysis. The sensitivity of the creep–fatigue damage summation locus for 1CrMoV rotor steel to different deformation and damage calculation solutions applied to the TMF and thermal fatigue test results is under evaluation in a similar way to that performed in the BE 5245 project [14].
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Fig. 6. Inspection at the 3D feature specimen: (a) Image from MPI inspection at cycle N = 70 with crack network at locations 1 and 2. (b) Typical crack appearance at locations 1 and 2 at low (left) and high (right) magnifications: pictures from the post-test inspection (N = 115 cycles).
5. Conclusions
References
The thermal fatigue behaviour of uniaxial and a range of component-like 1CrMoV rotor steel specimens under service-like loading conditions has been investigated. The results provide valuable information on the cyclic deformation behaviour and cycles to crack initiation characteristics of 1CrMoV rotor steel subject to service-like conditions. In addition, they serve to validate the effectiveness of a previously established constitutive model based on isothermal material property data. In addition, the following conclusions are also drawn. The cycles to crack initiation in isothermal and servicecycle thermo-mechanical fatigue tests on uniaxial and circumferentially notched round bar specimens can be rationalized when the cycles involve hold times of similar duration at the same temperature and a consistent crack initiation criterion is adopted (in this case, a discernable crack extension increment of 0.1 mm). The simple use of isothermal LCF with 1 h hold time endurance data cannot be used in isolation to predict cycles to thermal fatigue crack initiation, in particular in circumstances where stress relaxation is constrained by high triaxiality in long hold times (e.g. as in the 3D specimen thermal fatigue tests). In such cases, the prediction of component life requires a more comprehensive assessment of creep and fatigue damage.
[1] Timo DP. Designing turbine components for low-cycle fatigue. In: Proceedings of the international conference on thermal stresses and fatigue, Berkeley; 1969. p. 453–69. [2] Mayer K-H, Tremmel D. The thermal fatigue of components in steam power plant. In: Proceedings of the international dvm symposium on low cycle fatigue strength and elasto-plastic behaviour of materials, Stuttgart, 8/9 October 1979. p. 105–16. [3] Ha¨rkegard G. Designing steam turbines for transient loading. In: Larsson LH, editor. High temperature structural design, ESIS12. Mechanical Engineering Publications; 1992. p. 21–40. [4] Dawson RAT. Monitoring and control of thermal stress and component life expenditure in steam turbines. In: Proceedings of the international conference on modern power stations, AIM, Lie`ge, September 1989. [5] Bhongbhibhat S. Untersuchungen u¨ber das Werkstoffverhalten in Gebiet der Zeitfestigkeit zur Erstellung von Berechnungsunterlagen fu¨r Uberwiegend Thermisch Beanspruchte Bauteille. Technischerwissenshaflicher Bericht, MPA Stuttgart, Heft 79-02, 1979. [6] Thomas G, Dawson RAT. The effect of dwell period and cycle type on high strain fatigue properties of 1CrMoV rotor forgings at 500–550 C. In: Proceedings of the international mechanical engineering conference on engineering aspects of creep, Sheffield, 1980. Paper C335/80. [7] Miller DA, Priest RH, Ellison EG. A review of material response and life prediction techniques under fatigue–creep loading conditions. High Temp Mater Process 1984;6(3, 4):155–94. [8] Bicego V, Fossati C, Ragazzoni S. Low cycle fatigue characterisation of a HP-IP steam turbine rotor. Low cycle fatigue. ASTM STP, vol. 942. West Conshohocken, PA: ASTM International; 1988. p. 1237–60. [9] Colombo F, Masserey B, Mazza E, Holdsworth SR. Simple modelling of the constitutive behaviour of a 1%CrMoV rotor steel in service-like thermo-mechanical fatigue tests. Mater High Temp 2002;19(4):223–5. [10] Masserey B, Colombo F, Mazza E, Holdsworth SR. Factors influencing the service-like thermomechanical fatigue test cycle
Acknowledgment The financial support of Alstom Power, Steam Turbines R&D, is gratefully acknowledged.
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endurance of 1%CrMoV rotor steel. Fatigue Fract Eng Mater Struct 2003;26(11):1041–52. [11] Mazza E, Holdsworth SR, Skelton RP. Characterization of the creep–fatigue behaviour of a 1CrMoV turbine steel. Mater High Temp 2004;21(3):119–28. [12] Mazza E, Hollenstein M, Holdsworth SR, Skelton RP. Notched specimens thermo-mechanical fatigue of a 1CrMoV turbine steel. Nucl Eng Design 2004;234:11–24.
[13] Holdsworth SR, Mazza E, Jung A. The response of 1CrMoV rotor steel to service-cycle thermo-mechanical fatigue testing. J Test Eval 2004;32(4):255–61. [14] Brite-Euram project BE 5245. Optimisation of methodologies to predict crack initiation and early growth in components under complex creep–fatigue loading (C-FAT). In: Holdsworth, editor. Synthesis report for publication. 1997.