2008 Experience With Rotor-Stator Interactions in High Head Francis Runner

IAHR 24th Symposium on Hydraulic Machinery and Systems OCTOBER 27-31, FOZ DO IGUASSU RESERVED TO IAHR

EXPERIENCE WITH ROTOR-STATOR INTERACTIONS IN HIGH HEAD FRANCIS RUNNER

ANDRÉ COUTU Andritz VATECH-HYDRO Ltd , 6100 Trans Canada highway, Pointe-Claire QC, H9R 1B9, Canada

MICHEL D. ROY

Hydro-Québec, 845 Ste-Catherine est, Montreal, QC, H2l 4P5 Canada

CHRISTINE MONETTE Andritz VATECH-HYDRO Ltd , 6100 Trans Canada highway, Pointe-Claire QC, H9R 1B9, Canada

BERND NENNEMANN

Andritz VATECH-HYDRO Ltd , 6100 Trans Canada highway, Pointe-Claire QC, H9R 1B9, Canada

ABSTRACT

Built for Hydro-Québec, unit 1 of Ste-Marguerite 3 (SM-3) was commissioned in April 2003. After few days of generation at peak efficiency, noticeable noise was heard from the unit. The machine was stopped for inspection after 200 hours and major cracks at outflow edge junctions to crown and band were discovered on many blades. A Root Cause Analysis (RCA) process was launched following this discovery.

The main reasons for the SM-3 failure were identified to be high dynamic stress due to the interaction of guide vanes and runner blades, acting at a frequency very close to the natural frequency of the runner. Although qualitatively known at that time, the quantitative importance of runner blade-guide vane interactions, also called Rotor-Stator Interactions (RSI), in the dynamic behaviour of the turbine has only been demonstrated recently. All runners in operation are excited by RSI. Although this excitation is most of the time very small and does not damage the runner, in the case of SM-3, the phenomenon was so strong, that it made the runner crack at an accelerated pace. The paper will describe the RSI phenomenon that took place in the SM-3 runner.

A temporary solution has been rapidly developed for SM-3 in order to put the unit back in service and minimize the loss of production. At the same time, using the new design rules defined during the RCA, a new runner was developed and tested. The hydraulic performances obtained exceeded the ones from the first runner while ensuring high level of protection against RSI. Strain gage measurements performed on the newly installed replacement runner confirmed the effectiveness of the approach taken to design the new runner.

Hydro-Québec’s initial specifications for SM-3 have been put together using their experience with the 318m head Churchill Falls runners, which have been running trouble-free for decades. Following SM-3 experience, these specifications have been updated to avoid such events for the replacement runners and future projects. KEY WORD: Francis turbine, fluid-structure interaction; rotor-stator interactions; flow-induced vibration,

24th Symposium on Hydraulic Machinery and Systems

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INTRODUCTION

The Sainte-Marguerite 3 (SM-3) project, located in the province of Quebec, 600 km north-east of Quebec city, houses two Francis units which have a rated output of 447 MW under a net head of 330 m at a rotational speed of 257,1 RPM. The contract for the two units was awarded to GE Hydro in 1996.

The hydraulic design was challenging due to the wide operating range: the peak efficiency was located at approximately 67% of rated power to maximize the weighted average efficiency. Except for this special requirement, hydraulic and mechanical designs were straightforward. Using the state-of-the-art design tools available at the time, everything indicated trouble free commissioning and commercial operation.

Unit #1 was put into commercial operation in April 2003. After less than two weeks of operation, a severe runner blade-cracking situation was discovered. Inspection showed cracks on several blades at outflow / crown and band junctions varying from few to approximately 35 centimeters as shown in Figures 1 and 2.

In agreement with Hydro-Quebec, a detailed RCA was launched to help identify the

source of the problem and to find a solution. The team included technical experts from across GE Energy and GE’s Global Research Center. The various actions taken in this RCA process, together with a detailed history of the events up to March 2004, have been presented previously [1]. Based on the latest developments, the present paper will emphasize on the causes of the failure. It will also show how solutions, temporary and permanent, have been implemented at SM-3 to solve the issue while minimizing down time of the units.

ROOT CAUSE

The first cracks were repaired and strain gages were installed on few blades to

measure stresses at the outflow edge junction with crown and band. As predicted by finite element analyses (FEA), the measured static stress was very low. What the strain gage test indicated, however, was the presence of unexpected very high dynamic stress on the blades. The measured stress was so high that the dynamic force required to produce it was on the order of magnitude of the one producing the power. The phenomenon, although very important, was undetectable from outside the runner. Another characteristic of the signal was the almost perfect sine wave at each strain gage (see Figure 3). The frequency of this signal matched the guide vane passing frequency and the phase shift from one blade to the other matched guide vane passing phase: The runner was failing under High Cycle Fatigue (HCF) coming from RSI induced dynamic stress.

Figure 1 Crack at outflow to crown junction

Figure 2 Crack at outflow to band junction


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Rotor Stator Interaction (RSI) The hydraulic effect leading to RSI is primarily a potential flow interaction between the

non-uniform flow distribution at the outlet of the guide vanes and the rotating runner blades passing through this flow. Since the flow field in the radial space between the guide vanes and runner is non-uniform circumferentially, both the static pressure and the velocity (velocity magnitude and flow angle) vary circumferentially. This results in three effects on the runner: • As each runner channel passes through the flow field, it is subject to a varying static

inflow pressure. If a phase difference between adjacent channels is present, this results in a varying pressure difference between those channels, i.e. an unsteady load on the blade;

• As each blade passes through the flow field, it is subject to a varying incidence angle. This in itself creates an unsteady load on the blade;

• As each blade passes through the flow field, the magnitude of the velocity passing over the blade varies. This also creates an unsteady load on the blade.

Figure 3 SM-3 - Strain gage measurements at a blade outflow edge junctions with crown and band

Figure 4 Rotor-Stator Interactions


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Figure 4 illustrates these contributing factors. Detailed explanations on RSI calculations using unsteady CFD have been presented previously [2]. A useful way of representing the effect of unsteadiness on the blades of a runner is the unsteady or dynamic blade torque. This is the torque an individual blade contributes to the shaft torque as a function of time. On SM-3 runner, it so happens that the peak-to-peak dynamic torque calculated was close to the static torque. Verifications performed later on other runners indicate that the SM-3 dynamic torque was many times higher than any of the reference projects available.

Natural Frequency

Although suspected during SM-3 RCA, the proximity of the guide vane passing frequency ω with the natural frequency of the runner ωn was later found to be a very important parameter contributing to the response of the structure to RSI [3]. If resonance takes place between one of the runner natural frequencies and the RSI frequency, the dynamic effects can be amplified to dangerous levels. It becomes therefore essential to accurately predict Francis runner natural frequencies in water.

Modal analysis of runners in air has been performed for many years. In order to evaluate natural frequencies in water, it was common practice to multiply natural frequencies calculated in air by an empirical factor to take into account the added mass effect of water. Today, however, it is possible to accurately predict Francis runner natural frequencies and mode shapes in water using computerized modal acoustic fluid-structure analysis [4].

Based on the number of guide vanes and runner blades in a turbine, only the modes of one specific nodal diameter can get excited due to RSI [3]. For SM-3, with 20 guide vanes and 15 runner blades, the Nodal Diameter (ND) of interest is ND=5. Calculations of the first ND=5 mode in water are presented in Figures 5 and 6. These computer tools being unavailable at the time of the SM-3 RCA process, full size natural frequency measurements have been performed on the shaft-runner assembly (Figure 7). The natural frequency measurements, confirmed later by the computerized calculations, indicate a ratio ω/ωn=0.96 for the first ND=5 mode. Such ratio, close to one, may lead to a substantial increase of the excitation, depending on the damping value, as shown in Figure 8 for a single degree of freedom system. The real importance of the damping and natural frequency on the response of the runner to RSI have been studied lately when the technology of forced response calculation in water became available [5]. It confirmed what was only suspected at the time of the SM-3 RCA.

Figure 5

ND=5 runner displacement mode shape

Figure 6 ND=5 water pressure mode shape


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Single degree of freedom system

amplification factor for various damping values

0

0.5

1

1.5

2

2.5

3

3.5

0 0.5 1 1.5 2 2.5 3

excitation frequency / natural frequency

am

plifi

ca

tio

n f

ac

tor

0 0.05 0.1 0.15 0.25 0.375 0.5 1

SOLUTION

Temporary Solution Late summer 2003, the conclusion of the RCA was that the runner was cracking due to

high cycle fatigue coming from RSI induced dynamic stress with some natural frequency amplification. The problem being defined, a solution had to be found and quickly implemented to minimize down time and water spilling.

A modification was designed for the runner inflow based on CFD results to decrease the RSI pressure fluctuations. The modification was implemented at site and the improvement was clear: The strain amplitude decreased by at least half, matching the dynamic pressure decrease predicted by the dynamic flow analysis. This improvement, although very important, was unfortunately not enough to bring down the stress to a harmless level.

A drastic decision was then taken: Struts will be added between the blades to bring down the dynamic stress to a safe level while a new runner would be designed and manufactured. In October 2003, the reinforced runner (Figure 9) was put in service. The stiffening up due to these struts had two effects: they decreased the dynamic stress, but they also increased the natural frequency of the runner, therefore moving away from resonance. It was measured that the ratio of the RSI frequency to the natural frequency was now ω/ωn=0.76. Comparison of strain gage measurements Fast Fourier Transform (FFT) for the original runner and the stiffened up one are presented in Figure 10. The temporary solution proved to be reliable: The runner was replaced by the new one only in October 2006, after three years of operation. In the mean time, the customer was able to reliably generate power and minimize water spilling.

Figure 7 Dip test of the SM-3 runner

Figure 8 Amplification factor due to resonance proximity


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Hydro-Québec New Requirements Hydro-Québec’s specifications for SM-3 have been put together using their experience

with the 318m head Churchill Falls runners, which have been running trouble-free for decades. The initial specifications for the runners where therefore limited to general requirements on fabrication made out of martensitic stainless steel (ASTM A-743 Grade CA6NM). Following SM-3 runner issues, these general specifications have been updated to avoid failures in the future. All new runners, including replacement ones to be supplied for SM-3, would follow these new specifications which cover the whole process of supplying a runner from design stage through all the manufacturing activities and testing at site. The most important additions are as follows:

At the design stage the specifications call for: • A static analysis by Finite Element (FE) demonstrating a maximum Von Mises stress of

200 MPa for normal operation and 360 MPa at runaway speed; • The mesh quality of the FE model has to be demonstrated according clearly defined rules.

One of them specifies that doubling the number of elements should not induce more than a 5% increase of the maximum calculated stress. Another one prescribes a maximum on the calculated error on the deformation energy;

• A fatigue analysis, taking into account the residual stresses and demonstrating a fatigue life of at least 70 years using a specified number of starts and stops, runaways and production levels;

• An harmonic analysis showing no risk of amplification of the dynamic loads;

Figure 10 Strain gage 1, strain amplitude FFTs of original and reinforced runner

Figure 9 Struts added to the runner


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• The level of high frequency dynamic loads has to be determined by a method accepted by Hydro-Québec and shall not exceed 60 MPa peak to peak to avoid high cycle fatigue failure.

New casting requirements were also implemented: • A limit of 0,035 % on the carbon content is imposed to improve the weldability; • On the first casted blade, 100% X-ray examination according to ASTM E446, E186 or E

280, depending of the thickness, is required to qualify the pouring process. The acceptance levels are strictly defined;

• On the crown and band, partial X-ray are required with the same acceptance criteria used for the blade;

• On all runner parts, 100 % Magnetic Particles (MT) according to ASTM E125 level 2 are required;

• Ultrasonic Tests (UT) on critical areas are required according to ASTM A609 level 1 for a dept of 0,005D and level 2 from 0,005D to 0,01D, D being the outlet diameter of the runner;

• The resilience of attached casting samples has to be at least 50 Joules at 0 ºC; • The hardness of repair welds, after heat treatment, shall not exceed 325 Brinell. The

hardness in the heat-affected zone shall not exceed 350 Brinell.

The new requirements also cover runner fabrication: • All parts have to be fully machined; • 100% of the blades surfaces and area of crown and band to be welded have to be verified

by MT according to ASTM E125 level 1; • The welding material have to be homogenous to the castings; • Welders have to be qualified according to ASME; • The welding process has to be tightly monitored; • The welding has to be done with a monitored preheat of 100 °C minimum and 150 °C

maximum. Preheat has to be applied on the whole crown or band and on the portion of the blade to be welded. Preheat should not be interrupted during the whole welding process;

• 100% of the welds have to be verified by UT according to ASME Pressure Vessel Code Division 1 Appendix 12. Critical areas have to be X-rayed according to article 12.5.4 of CSA-W59;

• The runner has to be post weld heat treated according to specific requirements. The temperatures during the process have to be monitored. The temperature rise should be such that the difference of temperature between the thickest and the thinnest part is never more than 50 ºC. The nominal heat treatment temperature (± 600 º C) should be maintained for a minimum of six hours;

• The cooling must be controlled down to 80 ºC in furnace; • After the heat treatment, the welds have to be verified by MT or penetrant (PT) according

to ASTM E125 level 1 or CCH-PT 70-3 class 3. Critical areas have to be verified by UT according to ASME pressure vessel code division 1 appendix 12;

• At this stage, major repairs have to be done with martensitic welding followed by a heat treatment. Only minor repairs on non-critical area can be done with austenitic welding without heat treatment. This implies that the manufacturer should perform all NDT required before the first heat treatment so no major defect are discovered after the first heat treatment and subsequent machining;

• Finally, the first runner delivered at site has to be equipped with strain gages to verify that the design criteria are met.


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Since the release of these new specifications, different manufacturers have supplied sixteen new runners to Hydro-Québec and they mechanically all perform according to expectations. Of course these requirements have a cost and an impact on fabrication cycle. Hydro-Québec considers however that these drawbacks are relatively minor compared to production losses due to failure of runners. For example, on a 200 MW turbine, the payback of the extra costs imposed by the new requirements represents only few days of production. Permanent Solution: The New Runner

In addition to Hydro-Québec’s new requirements, the new runner had to undergo the same approval process the original went through. Using the latest design tools, which now included RSI unsteady calculations, and the knowledge obtained from the first runner, a new one was designed. The dynamic torque was decreased by a factor of more than three, without compromising on performance. The new runner offered even better performance than the original one, as shown in Figure 11, and would probably have been chosen instead of the original one, had it been available at the bid time.

In parallel to the hydraulic characteristics of the runner, great care was also given to its

mechanical design. The dynamic behavior of the runner was closely looked at to minimize the dynamic stress at critical locations. Special attention was also given to the natural frequency of the runner in water to avoid possible resonance. All customer’s new requirements were met and the new runner was commissioned in October 2006. Strain gage measurements indicate that the dynamic stress level was not even a concern as shown in Figure 12. Both runners have now been replaced and the units are fully operational.

Figure 11 Comparison of the two SM3 runners


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FORCED RESPONSE CALCULATION Modern high head and high performance Francis runners are subject to complex

dynamic forces that can lead to high cycle fatigue and eventual cracking and failure of the turbine. SM-3 experience has revealed the effects of strong fluid-structure interaction on the natural frequencies, non-linear damping and the forced response. Since then, a major research effort has been underway to understand the fundamental characteristics of these machines in real operating conditions and to develop advanced physics based computational methods that can predict these characteristics [3]. Validation with field test data is part of this effort and the strain gage measurements of SM-3 have been very useful to calibrate the forced response calculations, which are now reality [5].

Figure 13 presents results of dynamic stress at outflow edge to crown junction calculated using the newly available forced response calculations in water. Figure 14 presents the comparison between the strain gage dynamic measurements performed on the original runner and the forced response calculated values. As one can see, the match is very good.

Figure 12 Strain gage 1, strain amplitude FFTs of original and new runner

8 µµµµεεεε

Figure 13 Von Mises dynamic stresses due to RSI


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CONCLUSION

SM-3 original turbine design was performed with the state-of-the-art tools that were available at that time. There was neither indication nor any design tools that could have been used at the design stage to predict the situation observed at SM-3. It was later discovered that RSI actually exists on all type of Francis turbines, with greater importance on the higher head range [6]. Although SM-3, with its 330m head, falls on the low end of this category, machines with much higher head never experienced such high failure rate. SM-3 was special: All factors that were later discovered to impact the dynamic behavior of a turbine happened, by coincidence, to have unfavorable values on the SM-3 design.

At this point, ANDRITZ VATECH HYDRO is in position to assess RSI dynamic phenomena at the design stage, therefore avoiding any new occurrence of similar situation. More developments are undergoing to further refine the analysis.

BIBLIOGRAPHICAL REFERENCES

[1] Coutu A., Proulx D., Coulson S., Demers A., Dynamic Assessment of Hydraulic

Turbines – High Head Francis, Hydrovision 2004, Montréal, QC, August 15-18 2004. [2] Nennemann B., Vu T.C., Farhat M., CFD prediction of unsteady wicket gate-runner

interaction in Francis turbines : A new standard hydraulic design procedure, Hydro 2005, October 17-20 2005, Villach, Austria

[3] Coutu A., Velagandula O., Nennemann B., Francis Runner Forced Response

Technology, Waterpower XIV, Austin, TX, July 19th 2005. [4] Monette C., Coutu A., Velagandula O., Francis Runner Natural Frequency and Mode

Shape Predictions, Waterpower XV, July 23-26 2007, Chattanooga TN, USA [5] Coutu A., Monette C., Velagandula O., Francis Runner Dynamic Stress Calculations,

Hydro 2007, Granada, Spain, October 15-17 2007 [6] Coutu A., Aunemo H., Badding B., Velagandula O., Dynamic Behaviour of High Head

Francis Turbines, Hydro 2005, Villach, Austria, October 17th, 2005

Figure 14 Comparison between measured dynamic strains and forced response

calculationsVon Mises dynamic stresses due to RSI

2008 Experience With Rotor-Stator Interactions in High Head Francis Runner

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