GT2009-60144-Psarra

1 Copyright © 2009 by ASME

Proceedings of ASME Turbo Expo 2009: Power for Land, Sea and Air GT2009

June 8-12, 2009, Orlando, Florida, USA

GT2009-60144

FINITE ELEMENT TURBINE BLADE TANGLING MODELLING FOLLOWING A SHAFT FAILURE

Aikaterini Psarra Cranfield University

School of Engineering Dept. of Power and Propulsion

Gas Turbine Engineering Group Cranfield, Bedfordshire

MK43 0AL, England [email protected]

Vassilios Pachidis Cranfield University




Pericles Pilidis Cranfield University




ABSTRACT

A shaft failure in a gas turbine engine is a severe event which leads to a sudden decoupling between the compressor and turbine, while there is not any instantaneous variation in the aerodynamic power flow. During a shaft failure event, the decoupled turbine is free to accelerate to a terminal speed whilst, depending on the arrangement of the shaft support bearings, the aerodynamic loads may also force it to move rearwards and contact the downstream NGV structure. If the terminal speed attained exceeds a certain critical limit, high energy debris may be released from the engine compromising the safety of the operations.

In order to prove that shaft failure events can be handled in a safe and contained manner, engine manufacturers need to demonstrate among others that the extremely high rotational speeds a free running turbine can attain, can be reduced to a minimum value as quickly as possible. The present paper attempts to prove that one potential mechanism for limiting terminal speed may be blade tangling. Seal segments and platforms in particular can be designed in such a way so that they become quickly damaged and eroded by the dislocated turbine’s disc to allow for a quick contact between the turbine rotor blades and NGVs. A premature blade tangling can increase the energy dissipated as friction and heat between the structures and can lead to a decrease in terminal speed.

The work reported here investigates this exact scenario focusing on a hypothetical intermediate pressure (IP) shaft failure of a modern 3-spool High By-pass Ratio (HBR) turbofan engine. The study investigates the effects of the various damage mechanisms considering the violent interaction of turbine structures using Finite Element Analysis

(FEA). More specifically, the paper discusses analytically the development of a three-dimensional FEA model for the simulation of the dynamic impact phenomenon as well as the implementation of a dynamic non-linear finite element solver for the modelling of blade to vane interactions. A number of sample scenarios involving IPT blade to LP1 vane contact are presented to provide a better understanding of the effects of blade tangling on the evolution of the event.

The study reported in this manuscript constitutes a first important step towards developing an appropriate simulation strategy for the modelling of turbine interactions following a shaft failure event. It seeks to advance today’s knowledge in the evolution of such complex events and the effects on turbine terminal speed of blade tangling and energy dissipated in eroding/melting surrounding structures.

NOMENCLATURE D Cowper-Symonds Coefficient FD Dynamic Friction Coefficient FS Static Friction Coefficient HBR High By-pass Ratio IPT Intermediate Pressure Turbine LP1 1st Row Low Pressure NGV Nozzle Guide Vanes p Cowper-Symonds Coefficient ε Plastic Strain µc Friction Coefficient σy Yield Stress σd Dynamic Yield Stress vrel Relative Velocity

Proceedings of ASME Turbo Expo 2009: Power for Land, Sea and Air GT2009

June 8-12, 2009, Orlando, Florida, USA

GT2009-60144


INTRODUCTION

The development of aero gas turbines demands a rigorous effort to prepare and integrate computational models [1,2] and hardware, generate analysis results and post process data [3,4] in order to access the operational safety of the machine and generally its airworthiness. Proving the mechanical integrity of an aero engine under various nominal and non-nominal operating conditions is crucial for satisfying durability and reliability requirements. Moreover, from an engine certification point of view, engine manufacturers need to demonstrate that a catastrophic failure of an engine component is highly unlikely and that in any case it will not compromise the safety of the aircraft and passengers. This includes shaft failure events.

A shaft failure is a fairly complex and potentially hazardous event. It can trigger a number of mechanisms which can have an effect on the mechanical integrity particularly of turbine discs and attached blades. Following a shaft failure, the free running turbine will over-speed within the first few milliseconds after the event due to the sudden decoupling between the compressor and turbine. If the terminal speed of the turbine exceeds the critical limit, the turbine structure will fail, releasing high energy debris. Depending also on the arrangement of the shaft support bearings, the free running turbine may become dislocated and move downstream causing a mechanical interaction between the rotor and the surrounding stationary structure (NGVs and seals). In this case, the frictional energy that is dissipated between the structures may have a critical effect on the evolution of the event. For example, in the case of an IP shaft failure, tangling of IPT blades with LP1 NGVs would cause a significant increase in the frictional energy, which would affect in consequence the maximum rotational speed attained by the IPT.

Figure 1 represents a schematic of the IPT and NGV structures and indicates the areas that are most likely to come into contact following a shaft failure. After the IP shaft failure, the IP turbine rotor decoupled from the compressor behaves as a free running turbine. The aerodynamic forces resulting mainly from the impingement of hot gases on the surface of the blades are responsible for the rearward movement of the turbine rotor and the contact interaction with the stationary structure of NGVs and cavity seals. Obviously, the most critical loads in this case are the axial loads acting on the turbine blades and disc, arising from the gas pressure differences in the main gas path and the secondary air system, as well as the centrifugal forces imposed on the turbine rotor due to the over-speed.

In this particular case, as it can be observed in Figure 1, the distance between the IPT lock-plate and the seal segment and platform is smaller than the distance between the blades and the vanes. This implies that the first material interaction will occur between the lock-plate of the rotor disc and the seal segment/platform of the NGVs. Due to the high

axial loads and torque applied on the IPT structure, the material interaction of the disc and the seal segment and platform is so severe, that after a few milliseconds the surfaces of all involved structures are expected to erode (melt away) leading to a further IPT dislocation. The evolution of this material damage depends also on the material properties. A more ductile alloy has the capability of showing higher plastic deformation. Because of this, it is more likely to be damaged quicker resulting in advance blade tangling limiting, in this way, the terminal speed, than an alloy of lower plastic deformation. After the structural damage of the seal segment and the vane platform, the blade tips of the free running turbine are going to come in contact with the NGVs’ tips leading to high load on both surfaces and ,as a consequence, material deformation and possible failure.

It becomes obvious that in order to fully understand and capture the dynamics of such a complex phenomenon, a multidisciplinary analysis is necessary combining engine aero-thermodynamic performance simulation and mechanical integrity/structural modelling. The main aim is to investigate the effect of blade tangling on the rotational speed in terms of dynamic mechanical impact modelling.

NGV platformIPT lock-plate

Seal segment

Figure 1. Schematic of the IPT and the surrounding static structure

The majority of the research efforts that have been

reported in the public domain up to now have used broadly, finite element-based methods to identify the mechanical behavior and structural integrity of turbine blades and discs. They tend to concentrate on the analysis of stress distribution and fatigue crack initiation to estimate the lifetime of the aero engine. Some researchers have also proposed enhanced structures in order to satisfy durability and reliability criteria.

More specifically, Hou, et al. [5] investigated blade fatigue failures by mechanical analyses utilizing non-linear FEA to determine the steady-state stresses and dynamic characteristics of the turbine blade. They also looked into the contact interaction between the blade and the disc fir-trees under service conditions, taking into consideration significant geometric features, centrifugal forces and temperature


distribution. In 2006, Wiket [6] published a work on the damage mechanisms of the turbine disc, at the region of the lower fir tree slot, subjected to both operational and over-speed conditions. Apart from several studies on fatigue crack propagation, more recent papers have reported the transient vibration and friction forces of shrouded blades. In 2003, Petrov [7] demonstrated a method for analyzing the periodic force response of non-linear symmetric structures using a finite element model of a shrouded turbine disc capable of accounting for friction forces and interferences at nodes at the surfaces of shroud contacts. In addition, Szwedowinz, et al. [8] created a non-linear dynamic model to simulate the contact stress and friction of the shroud connection in order to assess the reliability of shroud couplings, whereas Sang-Ho Lim, et al. [9] investigated structural dynamics problems on bladed discs due to blade mistuning.

A predominant aim of engine design and development has always been to prevent any rotor failure that could release high energy debris [10]. Due to the fact that experimental test arrangements are considered not to be time and cost affordable, scientists tried to apply FE tools to investigate the effect of a possible failure on aircraft safety. In this way, past studies have been focused on predicting the containment of disc burst fragments [11], the blade release [12] and evaluating the turbine clashing behavior after shaft failure [13].

Generally speaking, FE tools have proven to be an accurate and time efficient way to analyze the structural dynamics of turbine blades and discs taking into account stress distribution, friction and impact forces and vibration. More published works [14, 15, 16, 17, 18] have presented the implementation of FE tools for the study of other engine components such as fan blades and shafts. Several FE models and methods have been utilized up to now for the structural analysis of gas turbine engine components. The majority of the work being done however tends to focus on particular isolated phenomena and usually aims at improving the design of existing engine structural systems. To the authors’ best knowledge a simulation strategy for the analysis of complex failure scenarios that are associated with severe structural interactions has not been reported yet in the open literature.

This study utilizes three commercially available software tools: i) a high performance pre-processor, Altair Hypermesh, ii) a finite element program for non-linear response of 3D structures, LS-DYNA3D and iii) a dedicated post-processor for LS-DYNA3D, LS-POST.

HyperMesh is a high performance pre-processor which supports the LS-DYNA3D Finite Element package. It provides tools to build and edit models with its specific 2D and 3D mesh generation panels and comprehensive meshing capability. LS-DYNA3D is an explicit finite element program able to analyze the non-linear dynamic response of three-dimensional structures and gives the opportunity to solve complex crash problems like the one analyzed in this paper.

Finally, LS-POST is a dedicated post processor for LS-DYNA3D which visualizes the results obtaining time histories for nodes and elements.

The main objective of this study was to develop a simulation strategy that has the potential to investigate complex structural interactions and particularly blade to vane contact. This paper presents the simulation strategy developed for the analysis of the interaction between an IPT and LP1 NGVs in the case of an IP shaft failure scenario. Sample results of rotor blade to vane contact are presented together with a parametric analysis of the effects of various boundary conditions.

SIMULATION STRATEGY

The investigation of the mechanical structure interaction between turbines after a shaft failure can be extremely computationally expensive and therefore demands the creation of separate high-fidelity models to simulate various key aspects of the dynamic impact phenomena. The simulation strategy followed by this study effectively breaks the overall complex event down to two key areas of interest that can be studied separately. More specifically, the first part of the analysis includes the interaction between the disc/lock-plate and the seal segment/platform of the NGVs as it is illustrated in Figure 2. The lock-plate and seal segment are the first parts of the structure that come in contact after the dislocation of the turbine. The second part involves the interaction between the turbine blades and the downstream turbine stators. In order to save computational time and for the sake of setting up a preliminary 3D model, the initial analysis of blade to vane contact included the material interaction between only two rotor blades, arranged axi-symmetrically, and a nozzle guide vane. Figure 3 presents the configuration of a blade and a vane.

Figure 2. 3D structure of disc/lock-plate - seal segment/platform [19]

Lock-plate

Seal segment

NGVs’ Platform Disc


The 3D FE models presented in Figures 2 and 3 reflect the real geometry of a modern three-spool HBR engine. The exact dimensions of the real geometry are company proprietary information and, as a result, they cannot be included.

Figure 3. 3D structure of a rotor blade and a stator vane The structure interaction between the lock-plate of

the IPT rotor and the NGVs’ platform and seal segment results in eroding part of the area of the seal segment and the platform. A consequence of the rubbing of the NGVs’ structure is the movement of the IPT further rearwards causing the tips of the vanes and the blade to come in contact.

Although, the FEA tool has a lot of capabilities, an appropriate implementation of critical parameters is vital to derive the desirable solution. Parametric analyses can give an insight to the understanding of the contact interaction leading to a suitable simulation strategy that needs to be followed for the accurate modelling of turbine structure interaction events.

Figure 4 represents the steps taken by this study for the modelling of a complete turbine failure scenario. Due to the complexity of the event, the simulation methodology is split into two structural models (as described above) which can run in parallel, reducing this way the total amount of computational power and time needed. Parametric studies are then carried out separately, and results obtained are grouped together in tabulations in a meaningful and appropriate manner so that when combined together they can give a complete picture of the evolution of a shaft failure event (i.e. axial displacement, energy dissipation, rotational speed and torque change with time etc). The biggest advantage of the proposed methodology is that results obtained could be converted into a non-dimensional form and could be utilized later on for other engine geometries.

Figure 4. Simulation strategy flow chart

MODELLING APPROACH

Structure and Mesh

The analysis reported in this manuscript focuses mainly on the high relative velocity impact between the rotor blades of the IP turbine and the stator vane of the first row of the LP turbine. As mentioned previously, in order to save computational time and for the sake of setting up a preliminary 3D model, the initial analysis of blade to vane contact included the material interaction between only two rotor blades, arranged axi-symmetrically, and one nozzle guide vane. The configuration of the structural model is illustrated in Figure 5.

It is important to point out that the study focuses particularly on the interaction between the trailing edge of the rotor blades and the leading edge of the stator at the tips. In order to achieve a considerable reduction in computational time, the areas coming in contact are modeled as deformable bodies, while the rest of the structure is considered to be rigid. Each blade consists of 66,356 elements while the vane is constructed by 52,712 elements.

Develop a 3D model of disc/lock-plate - seal

segment/platform interaction

Conduct parametric analyses to evaluate the effect of critical parameters including temperature effects

Develop a 3D model of blade to vane contact

Group results into non-dim maps and develop generic tabulations

Study of complete turbine failure scenarios and apply the same strategy to other engine geometries

Structure interaction between IP Turbine and LP1 NGVs

Group results appropriately into tabulations and establish a generic simulation strategy


The geometry of blades and vanes is defined by several complex three dimensional curves. The creation of the mesh demands the division of blades into a number of volumes in order to produce a good quality mesh. Blades and vane are meshed with eight-node hexahedron and six-node pentahedron elements. All solid elements use one-node integration points that define a constant stress throughout the element. The mesh of both blades and vane has nearly the same element size to avoid any penetrations (Figure 5).

During the simulations, the combination of the high velocity impact, the existence of large forces for a long time and the one-point integration elements used, was found to lead to zero energy modes. These zero energy modes, called hourglass modes, provoke ‘instabilities’ affecting the nodes’ displacement. This behavior of the elements is related to the increase in hourglass energy. The primary location of the hourglassing is in the highly stressed regions of the blade tips. A beyond limits hourglass energy gives a doubtful outcome in terms of plastic deformation of the structures. However, by employing an appropriate hourglass control card, the use of one-point integration elements proved to be an effective solution to the hourglass energy problem. When an hourglass control card is employed, an additional hourglass stiffness matrix is defined reducing the produced hourglass energy. Therefore, in order to avoid the existence of high hourglass energy at the regions of contact, a strain co-rotational stiffness form for the 3D structure is assumed according to Belytschko-Bindemann [20] with low hourglass constant definition.

Material models

Appropriate materials for the construction of turbine blades and vanes are the nickel based alloys due to their exceptional high temperature strength. The nickel based alloy material assigned to the deformable regions is modeled utilizing piecewise linear plasticity model. This material type is a strain rate dependent material model. Instead of stress-strain curves for each different strain rate, it is possible to input a quasi-static stress strain curve, which will be scaled with the Cowper-Symonds coefficients. In order to investigate how the strain rate dependency affects the results, at the beginning only the “static” stress-strain curve is used. In a further step, values of Cowper-Symonds parameters impose the strain rate effects.

Analysis

Before the blade to vane contact takes place, the blades, as a part of the whole turbine, rotate at a high angular speed. Due to the centrifugal effects, the blades are deflected into a quasi-static shape. This preloaded state of the structure is calculated by LS-DYNA3D during an implicit static analysis which is the first part of the entire simulation. The blades are free to rotate about their axial direction, while the

center of rotation of the body is fixed and the rotational load is applied as a body force.

Figure 5. Meshing of deformable areas of structures

In order to study the material interaction between the blades and the vane, a number of assumptions are made considering the boundary conditions applied on the structure. In the described model, the applied load on the free turbine rotor due to pressure difference is defined as a dynamic axial displacement of the center of rotation. In shaft failure events, the dynamic displacement depends also on the dynamic impact phenomenon of the plastic deformation of the seal segment and NGV platform after the interaction with the lock-plate of the disc. In order to evaluate how the growth of the seal segment/platform plastic deformation influences the blade to vane interaction, an analysis of different dynamic displacement scenarios was carried out. This is discussed in detail in the next section.

The actual relative rotational speed between blade and vane is a key parameter in determining the total amount of frictional energy dissipated during blade tangling. A high relative rotational speed implies that significant plastic deformation is likely to occur in both structures following the impact. The accurate assessment of the plastic deformation of the materials involved during contact requires including a failure criterion for the material. The piecewise linear plasticity damage model is based on utilizing plastic strain as a damage mechanism of failure of the material. The implementation of the damage model demands a contact card able to remove failed elements. This demand leads to the choice of the *ERODING SURFACE TO SURFACE [21] contact algorithm.

Furthermore, the contact between the blades and the vane is simulated with a static friction coefficient of 0.7, a dynamic friction coefficient of 0.2 [22] and an exponential decay coefficient of 0.25. Equation 1 shows the dependency of the friction coefficient on the relative velocity between the structures. In cases of impact at a high relative blade velocity Vrel, the friction coefficient is defined by the value of the dynamic friction coefficient.


( ) relvDC

c eFDFSFD⋅−⋅−+=µ

Eq. 1

The structural model response is also subjected to

the torque which forces the rotor blades on a circular path around the mass center. Using as boundary conditions the initial angular speed and the torque versus time, the variation of the rotational speed is then calculated considering the moment of inertia of the structure and the energy dissipated due to impact. SIMULATION RESULTS

Axial Displacement Variation

In this study, instead of defining a load that is applied to the IPT structure, an axial translation is employed to demonstrate how the IPT blades will be displaced axially with time. The axial displacement is representative of the forces acting on the structure and affects the growth rate of material plastic deformation. The effects of the axial dislocation of the turbine rotor are illustrated in Figures 6 to 9 for three different cases of displacement versus time.

Figure 6 illustrates the three notional curves of structure displacement versus time that have been examined in this study. The curves represent a possible moderate axial displacement of the rotor blades. The erosion rate of the elements of the seal segment and the platform define when the first contact interaction between the blade and vane tips is going to occur. The axial translation is connected to the plastic growth propagation of the seal segment and platform.

0

0.0005

0.001

0.0015

0.002

0.0025

0.003

0.0035

0.004

0.0045

0 0.005 0.01 0.015 0.02 0.025 0.03 0.035 0.04 0.045

time [sec]

axia

l d

isp

lacem

en

t [m

]

case 1

case 2

case 3

Figure 6. Imposed axial displacement versus time

All simulations were initiated at a typical free-running turbine angular speed of 1022 rad/sec before blade to vane contact. The applied torque forces the rotor blades to rotate around the mass centre leading to one rotor revolution in approximately 6 msec. During the high speed impact against the turbine stator, the blade loses its initial velocity and becomes highly deformed.

1010

1012

1014

1016

1018

1020

1022

1024

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035 0.040 0.045

time[sec]

Ro

tati

on

al

Sp

eed

[ra

d/s

ec]

case 1

case 2

case 3

Figure 7. Rotational speed versus time

Figure 7 represents the variation of the angular speed with respect to time. The rotational speed starts decreasing as soon as the high velocity impact occurs due to the frictional forces arising on both structures. In a case of a moderate crash impact as in case 3, where the rotational speed remains almost unaffected, frictional forces have minimum effect due to the smaller contact surface area. A more intense impact, leading to a higher axial displacement in the same time (case 2), results in a larger decrease in the rotational speed due to a larger frictional energy dissipation.

0.00

5.00

10.00

15.00

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25.00

30.00

35.00

40.00

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035 0.040 0.045

time [sec]

Slid

ing

en

erg

y [

J]

case 1

case 2

case 3

Figure 8. Sliding energy of the contacting surfaces versus time

In Figure 8, each step increase illustrates the contact between one blade and the vane at each time. Due to the existence of only two blades in the model, after a step increase in sliding energy, a period of no material interaction follows until the axi-symmetric blade contacts the vane structure again. In case 1, when the rate of the plastic deformation of the seal segment is high, the energy dissipated between the surfaces increases immediately after the contact occurs. After 40 msec, in case 1 and 2, the turbine blades travel downstream covering the same distance, 2 mm. However, the rotational speed in case 1 is lower than in case 2 and also the sliding energy is 5 J higher. Based on Figures 6


and 7, it can be concluded that bigger axial displacements of the IPT lead to lower rotational speeds.

0.278350

0.278400

0.278450

0.278500

0.278550

0.278600

0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035 0.040 0.045

time [sec]

ma

ss

[k

g]

case 1

case 2

case 3

Figure 9. Mass change due to elements erosion

Furthermore, according to Figure 7, the rotational speed of case 2 decreases more rapidly towards the end of the simulation to nearly 1016.5 rad/sec, while the sliding energy is lower than the one in case 1. The mass change due to the element erosion is depicted in Figure 9, where it can be observed that the mass status of case 1 and 2 is approximately the same for the last 3 msec. But in case 1, the element erosion takes place after the first 15 msec, resulting in a less rapid decrease in rotational speed due to larger changes in the turbine moment of inertia with time.

Variation of material characteristics

Material characteristics play a significant role in the development of structure interaction in cases of high velocity impact. In the previous analyses material interaction was studied without taking into account the strain rate dependency of the material. Therefore, a further analysis looked into the effect of the material model behaviour on the results of blade to vane contact. A material option with no strain rate dependency was compared with an option of high dependency. When the “static” stress-strain curve is taken into consideration, permanent deformations tend to be higher than those expected. This occurs because some materials tend to increase their strength under dynamic load and high strain rate ( ε/dtε d=& ).

The sensitivity of the material to strain rate is defined using the Cowper-Symonds relation [20], which correlates the yield static stress to the yield stress obtained during a dynamic load with high strain rate:

p/1

y

d

D

ε1

σσ

+=&

Eq. 2

LS-DYNA KEYWORD DECKBY LS-PREPOST Time=0.045Contours of Effective Plastic Strain

LS-DYNA KEYWORD DECK BY LS-PREPOSTTime=0.045Contours of Effective Plastic Strain

LS-DYNA KEYWORD DECK BY LS-PREPOSTTime=0.045Contours of Effective Plastic Strain

Figure 10. Effective plastic strain and element erosion calculated without a strain rate dependency of the material

Figure 10 illustrates the effective plastic strain calculated with a material of a “static” stress strain curve and Figure 11 depicts the plastic deformation of a blade when a yield stress is calculated according to equation 2. While elements of the blades are eroded for a material option without strain rate dependency, effective plastic strain is not high enough to provoke elements’ erosion in case of a strain rate dependency of the material. This is obvious in Figure 11, where the maximum effective plastic strain reaches about 80% of the failure strain.

LS-DYNA KEYWORD DECKBY LS-PREPOSTTime=0.045Contours of Effective Plastic Strain

Figure 11. Effective plastic strain and element erosion calculated with a strain rate dependency of the material

The consideration of the strain rate effects in the simulation is a key factor as it affects the rotational speed. Figure 12 illustrates the variation of rotational speed with and without a strain rate dependency of the material while the same failure criterion and the same axial displacement curve are considered. The decrease of the rotational speed in case of a material option without strain rate dependency is large in

Nikola Kafedzhiyski


comparison with the variation of speed when sensitivity of the material to strain rate is defined according to Cowper and Symonds relation. This result shows that the energy dissipated between the structures in case without strain rate dependent material is higher than that of a strain rate dependent material. The small peaks on the rotational speed curve of the ‘static’ strain rate material are related to the decrease in mass of the blade which leads to a reduction of moment of inertia and an instant acceleration.

1017.0

1017.5

1018.0

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1019.0

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0.000 0.005 0.010 0.015 0.020 0.025 0.030 0.035 0.040 0.045

time[sec]

Ro

tati

on

al

Sp

ee

d [

rad

/sec]

without strain ratedependency

with strain ratedependency

Figure 12. Rotational speed versus time CONCLUSIONS

As mentioned previously in the main body of the manuscript, FE models and methods have been used routinely for the structural analysis of gas turbine engine components such as shafts, blades and discs. Most of the work reported in the public domain however, tends to focus on particular isolated phenomena and usually aims at improving the reliability of existing engine structural systems. A simulation strategy for the analysis of really complex failure scenarios, associated with severe structural interactions, has not been reported up to now in the open literature.

This study investigates the mechanical structure interaction between the IPT and LP1 NGVs of a modern HBR turbofan engine after a shaft failure. The prohibitive computational power required for this type of analysis is addressed via the creation of separate high-fidelity models to simulate various key aspects of the dynamic impact phenomena. More specifically, the simulation strategy followed here effectively breaks the overall complex event down to two key areas of interest that can be studied separately. The first part includes the interaction between the disc/lock-plate and the seal segment/platform of the NGVs, while the second part involves the interaction between the turbine blades and the downstream turbine stators.

This manuscript reports analytically on the second part of the analysis evaluating the effect of blade tangling on the reduction of the terminal speed. Simulations were carried out studying the influence of axial displacement on the

change of rotational speed, sliding energy and mass change. The main output reported is that different axial displacements of the IPT in time, result in different amounts of frictional energy being dissipated and hence, different turbine rotational speeds. Moreover, the evolution of the material deformation was found to change when a strain rate material is defined instead of a ‘static’ strain rate. The study also showed that in order to obtain realistic plastic growth propagation and elements erosion the specification of appropriate material characteristics becomes vital.

The analysis presented here establishes key elements of an effective simulation strategy that can accurately capture the effects of structure interactions for the study of shaft failure scenarios.

ACKNOWLWGMENTS The authors would like to thank Mr. Steve Brown and Mr. Arthur Rowe of Rolls-Royce plc for their support to this project.

REFERENCES

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GT2009-60144-Psarra