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Engineering Failure Analysis 13 (2006) 1338–1350

Investigation of the failure of the L-0 blades

Z. Mazur *, A. Hernandez-Rossette, R. Garcıa-Illescas

Instituto de Investigaciones Electricas, Av. Reforma 113, Col. Palmira, 62490 Cuernavaca, Morelos, Mexico

Received 5 September 2005; accepted 24 October 2005Available online 14 February 2006

Abstract

A last stage (L-0) turbine blades failure was experienced at the 110 MW geothermal unit after 1 year of operation per-iod. This unit has two tandem-compound intermediate/low-pressure turbines (turbine A and turbine B) with 23 in./3600 rpm last-stage blades. There were flexible blades continuously coupled 360 degrees around the row by loose coversegment at the tip and loose sleeve and lug at the mid-span (pre-twist design). The failed blades were in the L-0 row ofthe LP turbine B connected to the generator. The visual examination indicated that the group of 12 L-0 blades of rotorB on the generator side was bent and another group of 5 blades at 140 degrees from the first damaged group was also bent.The cover segments were spread out from the damaged blades and had cracks. Laboratory evaluation of the cracking inthe cover segments indicates the failure mechanism to be high cycle fatigue (HCF), initiating at the cover segment holesouter fillet radius. The L-0 blades failure investigation was carried out. The investigation included a metallographic anal-ysis of the cracked cover segments and bent blades, Finite Element Method (FEM) stress and natural frequency analysis(of blades/cover segments), fracture mechanics and crack propagation analysis. This paper provides an overview of the L-0blades failure investigation, which led to the identification of the blades vibrations within the range 250–588 Hz induceddue to unstable flow excitation (stall flutter) as the primary contribution to the observed failure.� 2005 Elsevier Ltd. All rights reserved.

Keywords: Failure analysis; High cycle fatigue; Steam turbine failures; L-0 blade failure; Metallurgical examination

1. Introduction

The last stage blade (L-0 blade) is one of the most important contributors to the performance and reliabilityof the steam turbine. With the last stage blades typically producing 10% of the total unit output, and up to15% in some combined-cycle applications, improvements in last stage efficiency can significantly impact theoutput of the total unit. Retrofitting an older last stage design with a modern diaphragm and last stage bladecan typically improve heat rate and.output by up to 1% [1].

Life extension against erosion and reduction of vibration stress are important for improving the reliabilityand maintainability of steam turbines last stage blades. The use of a continuously-coupled connection struc-ture instead of the grouped blades has proved to significantly reduce the vibration stresses and is currently

1350-6307/$ - see front matter � 2005 Elsevier Ltd. All rights reserved.

doi:10.1016/j.engfailanal.2005.10.018

* Corresponding author. Tel.: +52 777 3623811; fax: +52 777 3623834.E-mail address: mazur@iie.org.mx (Z. Mazur).

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implemented in many retrofits of steam turbines [1–5]. The blades are continuously coupled 360 degreesaround the row by an integral or loose cover segment at the tip and loose sleeve and lug at the mid-span. Thisprovides excellent damping and reduces the dynamic response levels of the blades [1].

Modern turbomachines operate under very complex regimes where a mixture of subsonic, transonic andsupersonic regions coexists. The trend for improved large steam turbine design towards higher aerodynamicblade loading and smaller physical size attracts much attention to the aeroelastic behavior of blades in tur-bines. Flow-induced blade oscillations (flutter) of the turbine can lead to fatigue failures of a constructionand so they represent an important problem of reliability, safety, and operating cost. Aeroelasticity phenom-ena are characterized by the interaction of fluid and structural domains. The cascade flutter is characterized byaerodynamic interaction among oscillating blades in the blade row. Its importance can be understood from thefact that the unsteady aerodynamic force on blades is heavily dependent on the interblade phase angle. Fromthis standpoint neighbouring blade rows, e.g. a neighbouring rotor or stator will have a considerable influenceon the unsteady aerodynamic force because blade rows are closely placed in actual turbomachines. It is knownthat the flow-induced blades oscillations are pronounced mainly at the zone of the 90–100% of the bladelength. Also aeroelasticity phenomena can occur while turbine is operating at low load/low vacuum. Duringturbine operation at a low load/low vacuum, the L-0 blade vibration stresses are increasing abruptly. Thisincrease (peak) of the blade vibration is induced by unstable flow with oscillation of a shock wave near thethroat of the blade tip passage [1,2]. Operation at a low load like initial load-hold and other conditions shallbe determined by considering such phenomena.

In many cases by using either tip or mid-span shroud (cover) design, the blade structural damping can beincreased enough to prevent blade flutter. However, the shrouded rotor blade design will cause the blade modeshapes to be complex, and in some cases both bending and torsion mode components can be present at thesame time in a single mode [6]. Unfortunately, most existing analyses deal with uncoupled simple mode shapemodels, which cannot be used for the combined bending and torsion system mode analysis, and judgmentbased on past experience and experiments must be used to determine the acceptability of a shrouded rotorblade design [7–10]. On the other hand, prediction of the forced response of shrouded disc assemblies is stilla challenging engineering task because of unknown excitation loads and friction damping effects [11].Recently, a remarkable progress in transient flow calculations allowed the prediction of more realistic excita-tion forces acting on the rotating blades [12–17]. Nevertheless, due to some current uncertainties in transientflow calculation and forced response of shrouded disc assemblies, as was reported herein, some cases of steamturbine low pressure blades failures with a continuously-coupled connection structure have been recorded.

This paper provides an overview such a case of the L-0 blades failure investigation due to flow excitation(flutter) as the primary contribution to the observed failure.

2. Background

The blade under evaluation was the 23 in./3600 rpm last stage blade (L-0) of a 110 MW geothermal turbinewhich consists of two tandem-compound intermediate/low-pressure turbines with steam condition 1.1 MPag/182 �C/84 kg/s. During the unit last overhaul (January 2004), the original grouped rigid L-0 blades werereplaced by flexible blades continuously coupled 360 degrees around the row by loose cover segment at thetip and loose sleeve and lug at the mid-span. The evaluation of these failed new installed blades was carriedout after 1 year of blade operation period in base load mode. The blade is made of AISI 410 stainless steel.

During unit normal operation period the water feeding pump was tripped due to a mechanical problem andas a result the unit vacuum was dropped from 648 to 570 mm Hg and the load from nominal to 72 MW (65%of load) at an interval of 42 s from the last stable point. After the next 26 s the unit was tripped due to a con-denser low vacuum condition. After the turbine was restarted there were measured high vibrations, whichforced the unit to be shut down-to carry out an inspection and related maintenance. The turbine visual exam-ination revealed that the group of 12 L-0 blades from the generator side of rotor B, connected to the generator,was bent in the direction opposite to the rotor rotation and that another group of 4 blades of the same row at140 degrees from the first damaged group was also bent, as is shown in Figs. 1 and 2, respectively.

Fig. 3 shows a general view of the blade cover segments spread from the damaged blades. They presentdamage in the form of bending, rubbing, loss of material and cracks. The damaged loose lashing sleeves which

Fig. 1. The group of 12 blades, L-0 row from the generator side bended at the tip.

Fig. 2. The group of 4 blades, L-0 row from the generator side bended at the tip.

Fig. 3. Damaged loose cover segments.

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couple the blades in the mid-span are shown in Fig. 4. The sleeves were separated from the damaged blades,crushed and fractured in separate parts.

Fig. 5 shows the detail of the damage of cover segment No. 19; heavy deformation, rubbing, fracture andseparation of the pieces of material is apparent. In Fig. 6 is shown the crack initiation on cover segment No.12. The crack is localized on the outer filet radius of the wall of the cover segment hole. There were more coversegments with similar cracks.

Fig. 4. Damaged loose lashing sleeves.

Fig. 5. Detail of the damage of the cover segment No.19.

Fig. 6. Crack initiation on the cover segment No. 12.

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3. Metallurgical investigation of the L-0 blade

The metallurgical investigation of the failed L-0 blade was carried out, and included metallography, SEM(scanning electronic microscopy) fractography and chemical analysis. The microstructure of the blade airfoil(bended tip zone) is shown in Fig. 7. The microstructure consists of tempered homogenous martenzite, free offailures, typical for forged stainless steel according to specification AISI 410.

Fig. 8 shows the microstructure corresponding to cover segment number 12 made of Titanium Ti–6Al–4V.The microstructure represents bimodal distribution of volumetric phase ‘‘a prime’’ (�60% Vol.) within a lam-inar matrix of phase ‘‘a + b’’ (�40% Vol.). It is a typical microstructure of Titanium alloy Ti–6Al–4V aftersolution treatment and aging. The grain size is very small (fine grains) and measure 20 lm approximately.Also, a lot of small transgranular micro cracks can be seen as indicated by the arrows in Fig. 8b. The trans-granular crack propagation is typical of a fatigue failure mechanism.

Fractography evaluation was carried out on the exposed crack surface of cover segment No. 12 (see Fig. 6)using scanning electronic microscopy (SEM) to determine the origin of the fracture. Fig. 9 shows the different

Fig. 7. The microstructure of the L-0 blade airfoil made of AISI 410 stainless steel (bended tip).

Fig. 8. Microstructure corresponding to cover segment number 12 made of Titanium Ti–6Al–4V.

Fig. 9. Fracture initiation and propagation zones on cover segment No. 12.

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zones of the fracture propagation surface of cover segment No. 12. The presence of striations (fracture slidingplanes) which are characteristic for fatigue mechanism of fracture propagation, are noticeable.

Each striation represents one cycle of fatigue and by measuring the distance between striations it is possibleto determine the velocity of crack (fatigue) propagation. Average inter-striation distance was 3.73 lm in thefracture propagation zone and 1.70–1.97 lm in the fracture initiation zone. On the fracture surface no beachmarks were found. The presence of beach marks on the fracture surface commonly indicates that more eventsparticipated in fatigue propagation. The beach marks divide the fracture surface in zones of different rough-ness, which correspond to different events of fatigue. The absence of beach marks on the fracture surface ofcover segment means that only one event participated in fatigue propagation.

The chemical analysis of the deposits present on the fracture surface of cover segment No. 19 revealed somequantity of Ni, Cu, Co and S. The existence of significant clearances between the cover segment hole and theblade tenon facilitate accumulation of the deposits (oxides) which come from other sections of the turbine andcan promote corrosion.

4. Blade stress analysis

Using the Finite Element Method (FEM), centrifugal stresses, deformation and natural frequency of theindividual blade and coupled blades at 3600 rpm were calculated. The calculation also included the determi-nation of stresses at the cover segment. The results were evaluated considering the maximum stresses devel-oped at the blade and cover segment, maximum deformation and possible resonances. The first four modalforms of vibration of the individual L-0 blade are shown in Fig. 10.

The maximum deformation of the individual blade was 6.7 mm and occured at the tip of the blade, as isshown in Fig. 11.

Centrifugal stress distribution at the individual blade at 3600 rpm is shown in Fig. 12, The maximum stress,451.7 MPa occurred at the mid-span of the airfoil, close to the lashing sleeves at the blade pressure surface.

The natural frequency of coupled blades is represented in Fig. 13. The first natural frequency is 115.28 Hz,the second is 195.52 Hz, the third is 357.17 Hz and the fourth is 358.29 Hz. It can be seen that the third andfourth blade natural frequencies are very similar. Further, that the first and second frequencies are larger thanthose of the individual blade, whereas the third and fourth frequencies are smaller, respectively.

Fig. 10. Vibration modes of the individual blade at 3600 rpm.

Fig. 12. Centrifugal stress distribution at the individual blade at 3600 rpm.

Fig. 11. Individual blade maximum deformation due to centrifugal force.

First mode

Second mode

Third mode

Fourth mode

Fig. 13. The natural frequency of continuously coupled blades.

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The reported natural frequencies of the blades (individual and coupled blades) confirm that no mechanicalresonances of the blades structure exist.

The maximum deformation of the continuously coupled blades was 2 mm and corresponds to the thirdmode of vibration (Fig. 13). As was mentioned before a maximum deformation of 6.7 mm was registeredfor the free individual blade.

The centrifugal stress distribution at the blade airfoil pressure surface is shown in Fig. 14. The maximumstresses, of 391.5 MPa, were registered at the airfoil mid-span below the lashing sleeves and were much smallerthan in the case of free individual blade (451.7 MPa). In this view (Fig. 14) cover segments and lashing sleeveswere eliminated to see the whole blades. As may be evaluated from Fig. 15a, the maximum stress level,510 MPa, was encountered at the surface of the cover segment in the contact zone between the blade airfoiland the cover segment. The stress level at the hole of the cover segment in the contact zone between the bladetenon and the cover segment (fracture zone) was lower; 250 MPa, as is indicated by arrow in Fig. 15b. Thestresses at the lashing sleeves were very low and for this reason they are not shown here.

Considering a yield stress of 574 MPa for a blade made of AISI 410 stainless steel, a yield stress of 925 MPafor the cover segment material (Titanium Ti–6Al–4V), and the maximum calculated stresses presented previ-ously, the security factor for individual free standing blade is 1.27, for coupled blades it is 1.47 and for thecover segment it is 1.81.

Fig. 14. Centrifugal stress distribution at continuously coupled blades.

Fig. 15. Stress distribution at the cover segment.

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5. Fracture propagation analysis of the cover segment

A fracture propagation analysis of the cover segment was carried out, based on the metallographic inves-tigation findings and rules of fracture mechanics. Stable fatigue fracture propagation is governed by Paris Law[18] according to

dadN¼ CðDKÞm ðmm=cycleÞ ð1Þ

where dadN is the velocity of fracture propagation; da the crack size increment during one cycle of fatigue; dN

the number of fatigue cycles; C and m the Empiric constants; DK the fatigue stress intensity.The fatigue striations distances on the cover segment fracture surface correspond to the crack size incre-

ment, da, during one cycle of fatigue. These distances were measured and its values fall within the range1.7 lm at the fracture initiation zone to 4 lm at the fatigue fracture end zone. It was observed that the pre-dominant inter-striation distance was between 3 and 4 lm. To successfully separate the cover segment com-pletely from the blade, the maximum crack size should be more or less the same as a depth of the coversegment hole in which the blade tenon is installed (Da = 7 mm). The total time of the fatigue event was26 s, which correspond to unit low load, low vacuum operation (72 MW, 570 mm Hg). This time includesthe periods of fracture initiation and propagation, the time required for the deformation and spread offof the cover segment, and also the time required for the blade airfoils to deform. Considering, arbitrarily thatthe time for fracture propagation was a fraction of the total fatigue event time of 7 s, the possible blade exci-tation frequency that could lead to the cover segment fatigue failure was calculated from

f ¼ Da=Dtda=dN

½HZ�; ð2Þ

where f is the blade excitation frequency; Da = 7 mm the depth of the cover segment hole; Dt = 7 s the time ofthe fracture propagation; da = 1.7, 3 and 4 lm.

Using a rearranged form of the Paris Low the stress intensity factor, DK, was determined from

DK ¼ffiffiffiffiffiffiffiffiffiffiffiffidadN

1

Cm

r½MPa

pm�; ð3Þ

where C = 2.66e-12 and m = 4.23 determined using nCode [18].For three options of striations distance, da = 1.7, 3 and 4 lm, the corresponding stress intensity factors and

associated excitation frequencies were DK = 23.57 MPap

m and f = 588.23 Hz, DK = 26.96 MPap

m andf = 333.3 Hz, DK = 28.86 MPa

pm and f = 250 Hz, respectively. Further, considering these frequencies

and the same options of striation distance, the length of fracture was calculated using

Da ¼ dadN� f � t ½mm�. ð4Þ

It was found that for frequency f = 333.3 Hz and striations distance da = 3 lm the fracture length wasDa = 7.5 mm, which is very close to the real cover hole depth (Da = 7 mm) considered in this analysis. Thisfrequency (333.3 Hz) is very close to the natural frequencies of the third and fourth modes of vibration ofthe coupled blades (357.2 and 358.29 Hz, respectively). Because the striations distance 3 lm was found pre-dominantly on the cover segment fracture surface, it can be concluded that probably the third torsional modeX, which is typically related to the turbine operation with low load low vacuum, may be responsible for thecover segment fatigue failure and its separation from the blades.

6. Discussion of the results

The results of the metallographic examination of the failed L-0 blade and cover segment indicate that thecrack initiation and propagation on the cover segment was driven by a high cycle fatigue mechanism. Stria-tions characteristic to high cycle fatigue were found throughout the whole fracture surface of the cover seg-ment. Also, whole fracture surface presented small transgranular cracks, typically related to fatiguefracture mechanisms.

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Beach marks were not encountered on the cover segment fracture surface. The presence and number ofbeach marks commonly indicate how many fatigue events participate in fatigue propagation; the beachmarks divide the fracture surface in zones of different roughness that correspond to different fatigueevents. The lack of beach marks on the fracture surface means that only one fatigue event participatedin fracture propagation. The fatigue striations distance corresponds to the velocity of stable fracture prop-agation da/dN.

Considering the L-0 blades real operation period (1 year), it may be concluded that the mechanical res-onance of the blades does not contribute to blade failure. This conclusion also indicates that fatigue failureof the blades/cover segments was not originated during continuous operation under vibration stresses, butduring transition events. If the fatigue initiation and propagation were during continuous operation underresonance vibratory stresses, the blade failure would occur practically immediately (after a few hours ofoperation). Analyzing the unit’s operation history since the date in which the L-0 blades were installed (ret-rofit), only one period of unit operation with low load/low vacuum lasting 26 s approximately, wasdetected; this is congruent with metallographic findings on the cover segment fracture surface (lack ofbeach marks).

According to [2,19], the steam turbine operation with low load/low vacuum is inducing L-0 blade excitation(vibration) by unstable flow developing high vibratory stresses. Fig. 16 shows steam flow stream lines distri-bution at the L-0 stage during low load/low vacuum operation [19].

Due to reduced mass flow, the steam conditions are variable along the steam path; there are zones of dif-ferent pressure, radial flows, counter flows, flow recirculation (flow instabilities). These operation conditionsgenerate blade excitation forces (torsional vibrations X mode), which can lead to blade failures. Unit opera-tion with reduced mass flow is also causing a reduction of the flow velocity in the same degree. This results inchanges of the blade entry flow incidence angle (change of stage velocity triangle); the flow is entering into theL-0 blades with negative incidence angle (in this case it was 33–36� approximately) striking the suction surfaceof the blade airfoil and exciting the blades (stall flutter). The pressure fluctuation, flow recirculation and coun-terflows, in conjunction with the negative incidence angle flow striking on the blades, developed excessivevibratory stresses causing fatigue fracture and spread out of cover segments. In turn, it caused blade structuralloosening and drastically changed the blade vibration damping characteristics; the blades operate as it theywere free-standing. Furthermore, the forces developed by the steam flow and the vibratory stresses finallycaused the deformation of the blades. The deformation of the free-standing blades was facilitated by the mod-erate stress safety factor used to design the blades.

Fig. 16. Steam flow stream lines distribution at the L-0 stage during low load/low vacuum operation [19].

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7. Conclusions

On the basis of the analysis of the results of the L-0 blade/cover segment metallographic examination, unitoperational parameters, blade natural frequency and fracture mechanics it may be concluded that the L-0 bladefailure was originated at cover segments which were fractured and separated from the blades, causing looseningof the blades and changing drastically their vibration damping characteristics. The cover segments fracture ini-tiation and propagation was driven by a high cycle fatigue mechanism and was probably due to the combinedeffect of the unit operating in low load/low vacuum conditions, resulting in transient excitation of the blades.

Considering the real L-0 blades operation period (1 year) it may be concluded that the dynamic character-istics of the blades (mechanical resonances) did not contribute to blade failure. This means that the fatiguefracture of the cover segments was originated probably during transient events and not during continuous(stable) operation under vibratory stresses.

Analyzing the unit operational history and the cover segment fracture surface, only one transition event,e.g. unit trip due to operation with low load low vacuum, was found, which was associated with fatigue stri-ations on the cover segments fracture surface with lack of a beach marks.

The turbine operation with low load and low vacuum typically results in steam flow instabilities-pressurefluctuation and flow recirculation. Also, due to reduced steam flow/steam velocity, the flow is entering into thestage with negative incidence angle striking the suction surface of the blade airfoil. As a result, high vibratorystresses were developed in the blades structure, causing fatigue failure of the cover segments and their spreadout from the blades, loosening the blade’s structure damping vibration system and a consequently the bladesoperated as free-standing.

The force of the steam flow with negative incidence angle, together with vibratory stresses, generate exces-sive stresses and deformations in the free-standing flexible blades, as well as partial blockage of the steam pas-sage, in which flow recirculation and counter flows also participate. The free standing blade’s deformation wasfacilitated by the moderate stress safety factor applied at the design of the blades.

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