UNIVERSITÀ DI PISA

FACOLTÀ DI INGEGNERIA

DIPARTIMENTO DI INGEGNERIA AEROSPAZIALE

DOTTORATO DI RICERCA IN INGEGNERIA AEROSPAZIALE

XVIII CICLO

TESI DI DOTTORATO

CCONVENTIONAL AND IINNOVATIVE AAEROSPACE

JJOINING TTECHNOLOGIES

CURRICULUM

COSTRUZIONI AERONAUTICHE

TUTORE

PROF. ING. AGOSTINO LANCIOTTI

PRESIDENTE DEL CORSO DI DOTTORATO

PROF. ING. GUIDO BURESTI

DOTTORANDA

ING. CLAUDIA POLESE

ANNO ACCADEMICO 2005 – 2006

ABSTRACT I

ABSTRACT

The present PhD Thesis was developed in the framework of several Research Programs in which the Department of Aerospace Engineering at University of Pisa was directly involved, with the general aim to characterized the fatigue behavior of conventional and innovative joints typically employed in the production of aeronautical and aerospace structures. Correct predictions could obviously improve design of optimized structures, maintaining or increasing actual safety levels: but the separate or synergic effects of several parameters must be considered to correctly reproduce structure in-service behavior. The logic-flow adopted for each aspect analyzed in the present research had his origin on a consistent database of experimental data, ranging from general mechanical properties material characterization to strain gauge evaluations of associated stress fields, moving toward a more comprehensive description of the faced problem by means of Finite Elements simulations, to finally reach, by fracture mechanics codes, a generalized model. Attention was mainly focused on parameters characterized by a lack of information in literature or unsatisfactory prediction models. Among others aspects, the effect of Secondary Bending on conventional longitudinal fuselage joints was deeply investigated, bringing to a clear improvement of fracture mechanics software. Furthermore were analyzed different technologies capable of decidedly improve joint performance in a standard architecture. Processes as Interference Fit, Split Sleeve Cold Working, StressWave and ForceMate were demonstrated to significantly retard crack initiation and propagation in critical points, introducing beneficial stress fields. Besides currently great efforts are made by the aerospace International Community to investigate the possibility of replacing standard joining methodology by new technologies which are expected to provide considerable weight saving and cost reductions, incorporating integral structures in aircraft constructions. The more promising one, the “Friction Stir Welding”, was widely explored during the present activity in different aeronautical and aerospace applications, demonstrating to be a very robust and reliable process.

I ACKNOWLEDGMENTS I

ACKNOWLEDGMENTS

My deepest thanks to my Tutor, Prof. Ing. Agostino Lanciotti, who has been always the inexhaustible supporter of my research work, giving me continuously a chance to improve a new aspect of my knowledge. I wish to express my sincere gratitude to Prof. Ing. Luigi Lazzeri, who has been extraordinarily helpful with his kind technical and human support throughout all my work. Special thanks to Prof. Ing. Giorgio Cavallini for the interest demonstrated in all my activities and his amazing suggestions during all my stay at the Department. I wish to express my gratitude to all people that in so many different ways making possible the accomplishment of this work:

To Dr. Ing. Luisa Boni for her huge friendship, the invaluable support and the precious collaboration To Dr. Ing. Enrico Troiani, luckily our collaboration has gone over the professional relationship by consolidating a really important friendship

To the whole Staff of the Department of Aerospace Engineering: to the technicians of the Structural Laboratory: Mr. Mauro Romagnoli, for his invaluable help, the constant attention and his priceless technical suggestions, Mr. Claudio Lippi, Mr. Luca Lombardi and Mr. Fabio Pioli, for the funny camaraderie demonstrated during all the experimental activity to the technicians of the Electronic Laboratory: Ing. Massimo Pisani, Mr. Gabriele Adorni Fontana and Mr. Andrea Rossi. to all the Administrative Staff, Mrs. Fernanda Taddei, the real mother of the Department, Mrs. Martina Casali, Mrs. Roberta Razzi, Mrs. Nuccia Paganelli and Mrs. Franca Braccini to study here without all of you would not have been the same.

To all the student that were involved in the experimental and numerical activities of this research activity; I hope to helped them to grow up as good engineers as they helped me to improve both as teacher that as woman: Gregorio Nazzani and Emanuele Vincentini, Federico Tarabella, Guido Becchi, my first FSW student and Ing. Paolo Becchi that gave me the opportunities to know Liguria, Elisabetta Fattorini, a precious friend all over these years, Fabio Forlani and Gianantonio Toniolo, Roberto Marcalli, that reinforced my personality, Igor Azara, Vittorio Cipolla and PierPaolo Pergola, Daniele Fois, Daniele Caprai, Lorenzo Cabiddu for the funny moments in the FEM and Sardinia worlds, Nicola Pasquale, my most faithful student, Carmela Fierro the first nice girl I met in Pisa, Francesco Fruci, Chiara Battistelli, my favorite housemate, Paola Apruzzese, Federico Ferrari, the most cool FSW student, for his great work, Gianluca Magagnini, for his impertinent friendship, and Daniela, Biancamaria Felice and her really nice sisters,

II ACKNOWLEDGMENTS I

Francesca, Florinda and Mariachiara, and her incredibly kind Family, Alessio Iantorno, who persuaded me to visit Calabria, Emanuele Calvagna, Giuseppe Allegra, her sister Marilena and the nicest Sicilian Family I have ever met, Saverio Picerno, a man of substance, Lia Ruggeri for our great chattering moments and Enrico, Ivan Andreani and Vjola Ristori for their kindness, AndreaSebastian Daquino and Manuel D’Eusebio, Marco Dezi, Gabriella Perrone for her precious optimized friendship and Daniele, Filippo Nogarin, Daniele Dignani and Stefano Maestrini, Sergio Urso Sculco, a great friend in every situation, Luca Niccolini, Anita Catapano for her great young energy.

I also want to thank all people involved in the several research programs: at Alcatel Alenia Space Italia Mr. Roberto Laudato, for his great enthusiasm throughout all the FSW activity, Ing. Elena Bonadero, Ing. Secondo Ottaviano, Mr. Angelo Camassa and Mr. Walter Capitani, our NDI men, and Dr. Monica Bosco at EnginSoft, where I found the answers to so many questions, Ing. Roberto Gonella, my more enthusiastic sponsor, for the great opportunities that gave me, Ing. Roberto Merlo, for his precious suggestions, Ing. Francesco Franchini, my optimized men, Ing. Horia Horculescu, for the pleasant assistance, Miss Letizia Palanti and Mrs. Mirella Prestini for their kind and great efficiency. at AerMacchi Ing. Francesca Nigro and Ing. Sergio Peyronel at Agusta Ing. Ugo Mariani at EADS Ottobrun Ing. Frank Palm at the ISS, Italian Welding Institute, Dr. Ing. Michele Lanza at Genova University Prof.ssa Carla Gambaro of for her kind introduction to the FSW world at ABAQUS Italia Ing. Roberto Vitali, Ing. Alberto Mameli and Ing. Sguanci for their guidance in the complex ABAQUS dimension at Piaggio Aero Industries Ing. Enrico Ciompi and Alenia Aeronautica, Breda Costruzioni Ferroviarie, ATR Toulose, Perini Navi, EADS Matra Datavision.

I would like to gratefully acknowledge all people that supported my stay at the DLR in Cologne: Dr. Claudio Dalle Donne, for the great attention that he paid to me and to my work, and his brilliant wife Ozlem, Ing. Anne-Laure Lafly, my wonderful French best friend,

III ACKNOWLEDGMENTS I

Ing. Haiyan Chen, that open my eyes on the Chinese dimension, Dr. Tommaso Ghidini and his precious wife Barbara, Dr. Ulises Alfaro Mercado, Dr. Tamara Vougrin, Mr. and Mrs. Willms, the whole technical staff of DLR-WF laboratories, particularly, Mr. Karl-Heinz Trautmann, for his fundamental supervision of the work, Mr. Uwe Fuchs, Mr. Peter Mecke, Mr. Cristian Sick, Mr. Heinz Sauer, Mr. Herbert Hinderlich, Mr. Klaus Mull, for their patience and their valuable collaboration.

A sincere appreciation to all my Engineer and not Engineer Friends that spent so long and joyful time with me in Pisa: Nicola e Francesca, Maurizio, Giovanni, Alessandro, Francesco, Nicola, Guido, Francesco, Rosella, Carmela, Cristiano e Sara, Loredana e Mario.

Special thanks to my adoptive Pietra’s Family for their priceless support to this complex experience: Morello, for our long conversation on my little world problems, Carla who created my passion for yoga and a more spiritual life, Federica, who discovered my lyric world, Philippe Gouron, Gabriele Bonelli, Massimo and Luciano Tagliavini for our memorable goliardic dinners and Dr. Mauro Cerchiai for his big support in every field of knowledge.

My most profound thanks to my Family that supported me during this long years of study: my father Claudio, probably the man most proud of me, my mother Siriana and my sister Cecilia that correctly still affirm that I spent all my life in a strange world, my grandfather Siriano and my precious grandmother Liliana.

Huge thanks to My Francesco for being so incredibly understanding, patient, and most of all for cheerfully tolerating all the distractions of this work and for providing so many wonderful reasons to look up from my computer. You opened my eyes to the world.

IV INDEX I

INDEX

ABSTRACT ...... I

ACKNOWLEDGMENTS ...... II

INDEX ...... V

1. GENERAL INTRODUCTION ...... 1

2. CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES...... 3

2.1 INTRODUCTION ...... 3

2.2 EFFECT OF BENDING IN FATIGUE CRACK PROPAGATION OF THROUGH CRACKS...... 3

2.3 TECHNIQUES TO ENHANCE FATIGUE LIFE OF STRESSED HOLES ...... 7

2.3.1 INTRODUCTION ...... 7

2.3.2 EFFECTS OF “INTERFERENCE FIT” FASTENERS IN FATIGUE LIFE OF CENTRAL HOLES ...... 8

2.3.3 EFFECTS OF “COLD WORKING” PROCESSES ON FATIGUE LIFE OF OPEN HOLES...... 11

2.3.4 EFFECTS OF “FORCEMATE” IN A LUG-FORK JOINT: WING TO FUSELAGE CONNECTION OF AERMACCHI M-346...... 18

2.4 CONCLUSIONS...... 22

3. INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES ...... 23

3.1 INTRODUCTION ...... 23

3.2 THE “FRICTION STIR WELDING” PROCESS ...... 24

3.3 “FRICTION STIR WELDING” FOR AERONAUTICAL APPLICATIONS ...... 25 V INDEX I

3.3.1 FSW ON 6013 ALUMINUM ALLOY: IMPACT OF “DIRECT WATER COOLING” SYSTEM...... 25

3.3.2 FSW ON 2024 ALUMINUM ALLOY: EFFECT OF FLIGHT LOAD SPECTRUM ...... 30

3.4 “FRICTION STIR WELDING” FOR AEROSPACE APPLICATIONS...... 35

3.4.1 FSW ON “WELDALITE” ALLOY: THE FAST2 ALCATEL ALENIA SPACE ITALIA CRYOGENIC TANK ...... 35

3.5 CONCLUSIONS...... 41

4. SELECTED PAPERS...... 42

4.1 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES ...... 43

4.1.1 BENDING ...... 43

4.1.2 TECHNIQUES TO ENHANCE FATIGUE LIFE OF STRESSED HOLES...... 61

4.2 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES ..... 151

4.2.1 “FRICTION STIR WELDING” FOR AERONAUTICAL APPLICATIONS ...... 151

4.2.2 “FRICTION STIR WELDING” FOR AEROSPACE APPLICATIONS ...... 178

5. PUBLICATIONS ...... 231

6. TUTOR ACTIVITY: FINAL THESIS IN AEROSPACE ENGINEERING...... 236

VI GENERAL INTRODUCTION I 1. GENERAL INTRODUCTION

The perennial and ongoing challenge in airframe structural design is to provide an optimum solution simultaneously satisfying the competing requirements for safety, performance, and cost. The essential requirement for structural performance has challenged design safety factors more than in any other branch of engineering. In turn, this demands the use of high strength materials and relatively high design and operating stresses. Also, as commercial airplane usage and life spans have increased, so has the need for more robust, durable, and corrosion resistant structures. Consequently, airframe designers are always operating within exacting constraints, which leave essentially small margin for error. Success requires an intimate knowledge and understanding of the operating environment, of the structural and material behaviors and the proper tools for accurate predictions. A clearly established approach and well-defined criteria are essential to this success. Today this philosophy exist and is universally accepted to govern current design practices: the international FAR regulations requires in fact a “Damage Tolerance” certification for aeronautical primary structures. “Damage tolerance” is the ability to sustain operating loads (up to limit load) in the presence of unknown fatigue, corrosion, or accidental damage until such damage is detected through inspections, including non-destructive inspection (NDI), or safe malfunction, and then it is repaired. All primary flight-loaded structures must be designed to be damage tolerant. This requires that the structure have sufficient damage growth properties and detection characteristics so that if damage were to develop at single or multiple sites, normal specified airline inspections would ensure that the damage is found before it reduces the residual strength capability below the limit load. This requirement is essential to provide the most comprehensive assurance of continued safety throughout the service life of an airplane. The “Damage Tolerance” approach has significantly influenced structural design: an optimum balanced solution in which all requirements are met must include a thorough understanding of the operating environment and an accurate external and internal loads prediction capability, coupled with the availability of high-performance materials and effective productive technologies tailored for specific needs and refined techniques for predicting fatigue and fracture behavior, validated by comprehensive test programs. This hard design task is made more efficient today because of the experience of the past 50 years, which has brought about comprehensive database and new design analysis tools.

1 GENERAL INTRODUCTION I Tear downs of older airplanes and fatigue test articles and large-scale component tests represent fundamental data to relate design solutions and crack growth capabilities of the structures. Prediction methodologies have evolved simultaneously with the increases in knowledge and computing capabilities: in fact the finite element method (FEM) has become an essential instrument to perform a large variety of stress analyses. The correct prediction of the operating stress/strain levels in the detailed structural components has given the engineer tools to size structure accurately without having to add unnecessary conservatism, ensuring both safety and airplane performance. This refined methods for structural analysis allow to solve major, highly complex problems and to better understand the important mechanisms or parameters that influence structure damage tolerance, often eliminating expensive tests. So new design solutions with superior performance as well as unquestionable higher safety can be obtained. In this process of continuous optimization of structure efficiency new technology advances have a chief impact: improvements in high-speed machining, advances in laser beam welding, and development of friction stir welding all potentially lead to the replacement of built up structures, containing dozens of mechanically fastened parts with monolithic machined or integrated structures. The fewer hand fit operations bring to a strong reduction in the costly work involved in fabrication and assembly and the often hidden costs associated with errors and rework. But when applying monolithic and welded integrated structures, great care must be taken to ensure that durability, damage tolerance, and fail safety considerations are thoroughly understood and evaluated. So, despite great improvement obtained during the recent years on predicting complex structure fatigue behavior, several aspect must still be investigate to achieve a robust generalized prediction model.

2 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I 2. CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES

2.1 INTRODUCTION

Riveting is the most popular joining technique in aircraft structures and surely is going to maintain this position for many years. Riveting is used to realize joints in different aircraft areas, but recently the attention of researchers was particularly focused on fuselage riveted joints, also as a result of recent catastrophic failures of lap-splice joints of transport aircrafts. In fact, even if the fatigue behavior of rivet joints has been studied for more than fifty years, the nucleation and propagation of fatigue cracks in these elements are still not exactly predicted. Nowadays more sophisticated calculations can be performed to better understand the different aspects of the fatigue resistance of riveted joints, but focused tests still represent an indispensable benchmark. Several experimental investigations has been carried out, which, among others, demonstrated that by increasing and controlling the force applied during riveting, large improvements on fatigue life of riveted structures can be achieved. But more than a few factors contribute to the fatigue resistance of riveted joints and taking all these factors into account in prediction models is a very difficult task, far away, for the moment, from its definitive solution; also for this reason, the common approach to this problem, such as to other problems, is to investigate separately the different factors, hoping in the near future to be able to group them in a single model.

2.2 EFFECT OF BENDING IN FATIGUE CRACK PROPAGATION OF THROUGH CRACKS

Introduction

One of the less examined factors that influenced the fatigue life of typical riveted joints is the presence of a bending stress, superimposed on the membrane stress. The bending stress is not necessary to transmit the load in a joint, as, on the contrary, the membrane stress is (membrane stress is used here to indicate the stress, constant in the thickness of a plate, which is created when a axial load is applied to the plate).

3 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I The bending stress is created as a consequence of the inherent eccentricity of single lap or butt joints; the ratio between the bending and the membrane stresses is named "Secondary Bending", SB. Strain gauge measurements of SB in longitudinal fuselage lap splices have shown high values of this parameter, to the point that, even if the main load condition was the pressurization of the fuselage, compressive strains were recorded near lap splice joints. The bending stress increases the stress on the faying surfaces of riveted joints, so that small surface or corner cracks nucleate in this area where inspections are very difficult to perform. Subsequently, small cracks propagate in through thickness cracks that can be easily detected when they emerge out from the rivet heads. After this event, the residual life of the structure is only a small part of its whole life, so accurate calculations are necessary for a reliable estimation of safety margins and residual life of the cracked structure. However, only few investigations have been carried out on the behavior of through cracks subjected to combined traction and bending stresses, while surface cracks have been deeply investigated both from theoretical and experimental points of view.

Activity

The aim of the presented activity was the production of experimental results on fatigue crack propagation of through cracks under combined membrane and bending stresses, and the evaluation of the results on the basis of stress intensity factor solutions for a part-elliptical through crack. More than 30 fatigue crack propagation tests on plate specimens subjected to in phase membrane and bending stresses were performed. The bending stress was introduced in the specimens by mounting them eccentrically with respect to elastic grip plates. The dimensions of the specimens and grip plates were established on the basis of a preliminary calculation by the ANSYS FEM code, whose results were in good agreement with the experimental ones obtained by a strain gauge instrumented specimen. Tests were carried out at five different values of bending to membrane stress ratio, SB = 0 - 0,55 - 1,25 - 1,8 - 2,23. In order to evaluate crack front shape evolution during crack propagation, some tests were interrupted at different percentage of fatigue life and the specimens were statically broken to correctly measure the crack front geometries. As expected crack front shape is more and more oblique increasing the SB ratio and the evaluation of the corresponding stress intensity factor along crack front became more complex. The only bibliography reference on this subject is the Finite Element investigation performed by S. A. Fawaz on oblique thought crack characterized by different ratios of crack geometrical dimensions. The results obtained by means of the 3D Virtual Crack

4 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I Closure Technique are implemented on the commercial AFGROW software for crack propagation prediction. The code estimated fatigue life is really too conservative respect to experimental propagation data to be useful in an effective damage tolerance model. Criticisms of stress intensity factors calculations were overcome by new ABAQUS FEM computations performed on experimentally observed crack fronts. So passing to deeply analyzing the AFGROW “oblique trough crack” prediction model, it was observed that the code correctly implemented the Fawaz’s database, but it evaluated crack propagation only by means of two interpolation points on a virtual quarter elliptic crack front at the front and rear sides of the through crack. The two punctual values of the stress intensity factors gave rise to a markedly fast propagation at the rear point of the crack front, so obtained too less oblique crack fronts. The AFGROW code rapidly commuted the “oblique trough crack” model in the “trough crack” model, obtaining a shorted fatigue life. Another factor not considered in the propagation code was the relaxation of bending stress during crack propagation. By means of ABAQUS FEM non linear analyses was correctly evaluated, in each test configuration, the SB decrease due to the reduced flexural stiffness at different crack lengths. Taking into account all these results new prevision were performed. At the lower SB values was possible to reproduce the experimental results by evaluating stress intensity factors only on the basis of the pure membrane stress acting on a trough crack with the dimension of the anterior length of the oblique crack front to emphasize the effect of the bending stress. The fully comparable results obtained at the lower SB values confirmed that the two approximations introduced, straight crack front and sole membrane stress acting, tended to counterbalance. On the other hand at the higher SB values a crack propagation rate evaluated as previously described was strongly overestimated. The problem was solved using a mean stress intensity factor values that generated an incremental cracked area equal to the incremental cracked area evaluated by considering local stress intensity factor values. So the high gradient of the stress intensity factor at the back face of the plate not accelerated the propagation at the plate back side, because the crack growth was influenced also by the adjacent stress intensity factor values. The linearly decreasing of the SB values during crack propagations was also taken into account. The new corrections introduced to the basic model of “oblique through crack” allow to correctly evaluate fatigue life and crack shapes under combined traction and bending stresses. All the experimental and numerical results obtained are entirely exposed into the published papers [1] and [2], fully reported in Section 4.1.1.

5 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I Conclusions

This part of the activity allowed to elaborate a very simple and reliable model to predict crack propagation under a combined stress state of tension and bending, typical load condition in conventional fuselage lap splice joints. The suggested correction to a widespread commercial software has been recognized as a really good improvement in the Damage Tolerance of conventional structures. The evaluated model must obviously superpose to all the other effects present in a real joint.

6 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I 2.3 TECHNIQUES TO ENHANCE FATIGUE LIFE OF STRESSED HOLES

2.3.1 INTRODUCTION

The fatigue resistance of aircraft structures is directly connected with the fatigue resistance of structural joints. Focused only on mechanical joints a possible distinction can be made between permanent connections, which are usually realized by means of solid rivets, and threatened fastener or mechanical joints that can be dismounted. During recent years different technological processes has been developed to enhance the fatigue resistance of critical holes: they are mainly based on the introduction of a residual stress field around the hole that can reduce mean stress or stress amplitude of the applied fatigue spectrum. Among others, three similar processes had been analyzed in the present activity. Aeronautical industries accept as a standard practice Split Sleeve Cold ExpansionTM and Interference-Fit methods for riveted joints and ForceMate® bushing installation for open holes or structural connections as lug-fork. The investigated techniques are obviously different from the technological point of view but quite similar in the theoretical and FEM analysis aspects. The shared characteristic is the initial introduction of an oversized pin into the processed hole. In Interference-Fit process the pin (or rivet) is permanently installed; the interference level can be chosen in order to produce only elastic strains surround the hole. In this way the resulting elastic stresses field reduces the fatigue cycle stress amplitude. Instead in Split Sleeve Cold ExpansionTM process the interference level is high enough to produce plastic strains and the pin is subsequently extracted. So the residual stress field around the hole decreases the fatigue cycle mean stress. In ForceMate® process both the positive effects are included: the installation of an high interference bush produces also the hole cold working. Besides advantages associated to the compressive residual stresses field, the interference mounted bush avoids the direct contact between the hole and possible connection pin. In order to provide an effective design tool, a large research activity started at the Department of Aerospace Engineering of the University of Pisa in the framework of the European project ADMIRE and in cooperation with some Italian and American Aeronautical Industries. Different FEM models of these three techniques, characterized by increasing degrees of approximation of the real processes, were developed. Finite element data were compared with results of focused test series and strain gauges verifications on typical aeronautical materials. The results obtained were really helpful to investigate

7 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I process parameter influence in enhancing fatigue life and to formulate effective theoretical provisional methods.

2.3.2 EFFECTS OF “INTERFERENCE FIT” FASTENERS IN FATIGUE LIFE OF CENTRAL HOLES

Introduction

Many experimental tests have clearly demonstrated the beneficial effect of interference fit on metallic structure fatigue resistance. Despite this, although a lot of work has been carried out for other hole fatigue-life-enhancement systems, such as cold working, there is little information on the quantitative evaluation of fatigue resistance of interference fitted holes. The problem is not simple: even if cyclic stresses in well designed fastener joints do not exceed, under normal load conditions, the yield stress of the material, interference can produce local plastic deformations. Interface conditions, such as actual contact zone and related properties, are usually not clearly defined. Besides when an external load is applied, both the rivet and hole deform; slippage and loss of contact can take place. Moreover, the evaluation of the complex multi-axial state of stress around an interference-fit pin is not an easy task. Taking all these factors into account in prediction models is not simple, unless by introducing strong simplifying hypotheses. So, as explained previously, the beneficial effect of an interference fit is well known, but is difficult to develop a reliable quantitative prediction model.

Activity

During the experimental activity of the European Project ADMIRE (Advanced Design concepts and Maintenance by Integrated Risk Evaluation for aerostructures) an increased fatigue life was observed for higher squeezed riveted joints, with a D/d ratio between the diameter D of the formed head and the initial diameter d of the rivet shank higher than 1.2, i.e. 1.4 and 1.6, [3]. Actually when a solid rivet is driven the second head is formed but the rivet shank is also plastically deformed so the hole is filled creating a beneficial interference state. In order to evaluate the stress field in the joint was firstly performed a complete FEM analysis modeling the real process. But this simulation immediately appeared quite

8 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I complex due to the several parameters to take into account during the analysis. In fact was quite difficult to introduce the correct material elastic-plastic behavior in each process phase and to exactly reproduce the different rivet forming processes under force or displacement control. In addition quite big solution differences were obtained if the remeshing FEM code capability was or not activated. On the other hand few check parameters were disposable for evaluating simulation quality. It was possible only to measure the final dimension of the former head and the strength or the position of the driven head machine during the process. The enhancement of fatigue life observed at the different squeeze ratios was obviously connected with the reduced stress amplitudes experienced by riveted holes during fatigue tests, but, unfortunately, the real behavior seemed to be still quite complicate to be correctly reproduced by an analytical program. In the real situation the rivet stem was highly deformed so the contact wasn’t constant around the hole and the interference level wasn’t the same along the stem and was not exactly the same in all nominally identical riveted holes. To clearly understood the phenomenon and elaborate a good design instrument a simpler specimen configuration was examined. Simple dog bone specimens, 1.6 mm thick in the aeronautical aluminum alloy 2024-T3, having a central reamed hole, 4 mm diameter, were filled with different steel pins in order to obtain 5 different interference levels in the range from 0.5 to 2.5%. The pin was the same for each interference levels in order to avoid possible scatter in the results connected to the tolerance on pin machining and the hole was protected from possible damage during the pin insertion by an FTI self-lubricated split sleeve so no wear was observed in the pins. Several tests were carried out to highlight the effect of the interference levels on the specimen fatigue lives. Besides two different FEM models were elaborated: the most complex one simulated the real insertion of the pin and the other one simulated only the fit interference by means of the FEM ABAQUS code “clearance” command. Stresses obtained in the critical diametrical area along a radial path were fully comparable so the simplest “clearance” model was preferred, being less time consuming. FEM results, in which the experimentally measured stress-strain curve of the base material was inserted, were confirmed by experimental strain gauges measurements on a scaled bigger specimen. So the same calculation was performed for fatigue tests specimen. Seven interference levels were examined from 0 to 4% and different applied stresses covering the range of stresses applied in fatigue tests. In each model three different steps were examined: pin insertion, application of the maximum stress and unloading to the minimum stress, R=0,1.

9 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I Circumferential strains obviously increase with higher interference values, instead circumferential stress after an 1% interference level starts to decrease due to material plasticity. But the Interference Fit produces a strong biaxial stress state composed, in the critical section, of a longitudinal stress, whose variation constitutes the actual fatiguing stress, and a transverse compressive stress, whose value, at a fixed interference degree, is quite constant in fatigue cycles. Moreover, after the pin insertion and the first fatigue cycle, subsequent fatigue cycles, at the same stress level, only cause elastic stresses. So hoop strain seams to be the simpler and representative parameter that can be converted into an equivalent uni-axial stress, whose minimum and maximum values in the fatigue cycle are given by (Seq)min = (Sy)min − ν (Sx)min and (Seq)max = (Sy)max − ν

(Sx)max. The direct comparison of the fatigue results can be made only correcting the effective stress ratio of each tests to the equivalent maximum stress in the case R = 0,1. The method correlates very well the fatigue test results obtained under different interference levels, which are almost reduced to only one fatigue curve. But some additional factors are present, such as fretting and the position of the critical locations: the examination of the fractured specimens shows, besides a large scatter in the results, that the critical location moves around the hole by increasing the interference level up to 40° – 50°. So, for any criterion adopted, calculations would have to be carried out in several locations around the hole, looking for the critical value of the parameter. For comparison some tests were carried out with faceted pins in order to realize a uni-axial stress state at the notch root: results demonstrated that the suppression of the transverse stress increases fatigue life and what's more the fatigue critical location is again the notch root. In this case the equivalent uni-axial stress evaluated by FEM models is able to correlate experimental results quite well, but positioned slightly over the open hole fatigue results. Moving toward the simulation of more realistic problems also a load transfer specimen was experimentally tested for three different interference levels. The specimen was realized by an interference pin insertion with an FTI sleeve interposition as previously done. The pin was not tightened in order to avoid load transfer by friction. In this way the notch root is in the more critical condition. The experimental results put in evidence the high beneficial effects of interference fit for this loading condition: in fact an interference of 1.25% increase the design stress at 106 cycles life of 22% for a no load transfer joint but of the 82 % for a load transfer joint. Also for this test a FEM model were realized in order to evaluate the stresses introduced by pin insertion and subsequently loading and unloading phases.

10 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I Results interpreted in terms of equivalent stress practically became a single line, but cannot be directly derived from the base material curve. Fracture surfaces observed by means of microscope clearly shown as previously an increased angle of preference fracture at higher interference levels. This point so needs to be deeply investigated. All the experimental and numerical results obtained are entirely exposed into the published papers [4], [5] and [6], fully reported in Section 4.1.2.

Conclusions

Fatigue tests were carried out on 2024-T351, thickness 1.6 mm, central hole specimens containing pins mounted with different interference-fit levels. Tests clearly demonstrated the beneficial effect of interference fit on fatigue resistance up to the maximum value examined, equal to 2.5%. The finite element method was used to evaluate the stress and strain fields around an hole after the insertion of the pins and subsequent fatigue cycling. Hoop strain at the notch root was successfully used as a parameter to evaluate an equivalent stress; the fatigue results obtained under different interference-fit levels practically reduced to only one fatigue curve when interpreted in this way, even if it was not possible to derive this curve from the fatigue curve of the base material.

2.3.3 EFFECTS OF “COLD WORKING” PROCESSES ON FATIGUE LIFE OF OPEN HOLES

Introduction

The process usually known as “cold working” has become a wide spread practice over the last 20-30 years, mainly used for in-service aircraft as a means of recovery from manifested fatigue problems or, less commonly, used for new aircraft to guarantee that the operational and economic lives of critical components meet or exceed the required design life. The cold working principle consists in producing in the fatigue critical region a self-equilibrating residual stress field, of compressive type in the most critical point, which will superimpose the tensile stress introduced by fatigue loads. This results in a lower mean stress value, with evident benefits from the point of view of the fatigue resistance.

11 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I While a number of cold expansion processes of fastener holes in metallic aircraft structure have been developed over the years, at industrial level, the “Split Sleeve Cold Expansion” process, developed by Boeing Company and distributed by Fatigue Technology Inc. (FTI), has been by far the most successful. This process is based on placing a longitudinally split sleeve within the hole and drawing an oversized tapered mandrel through this assembly. The hole is expanded to an extent sufficient to cause permanent plastic deformation, so that, upon the removal of the mandrel, the surrounding elastic material attempts to recover its undeformed state, producing a self-equilibrating residual stress field. In particular, the circumferential residual stress is of compressive type in an annular region adjacent to the hole. During the process the hole is protected by the sleeve which avoids the direct sliding of the mandrel on the hole and, consequently, which precludes hole damage that can trigger off fatigue failures. Furthermore, the sleeve has the function of limiting the axial displacement of the material, that, otherwise, might be subjected to a like an extrusion process. In order to reduce the mandrel pull force, the sleeve is lubricated and longitudinally split; the region at the edge of the hole subjected to a non-uniform radial expansion due to the split of the sleeve requires a final hole reaming. A number of studies on the cold working technique have been performed by simulating the process as only an uniform radial expansion of the hole. But all these models provide an axial-symmetrical solution that cannot account for the complex, three dimensional residual stress field. The “Split Sleeve Cold Working” is the most known system to produce compressive residual stress field, but it is not the only one method. TM StressWave is a relatively new method, developed by StressWave Inc., completely different from the previous one, which introduces compressive residual stresses prior to machining the hole, without any pre- or extra post-processing operations. The method, uses hardened indenters to produce plastic deformation in an area smaller than diameter of the hole that will be subsequently drilled.

Activity

The problem of the quantitative assessment of compressive residual stresses around a Split Sleeve Cold Expanded hole was initially faced in the present work by several FEM analyses characterized by different degrees of approximation of the real process. The FTI specifications for the cold expansion process can be divided into three main steps: hole expansion by pulling a tapered mandrel through the hole, hole recovery once the mandrel is removed and a finishing ream that removes a thin layer of material around the hole. The differences among the analyses performed by the ABAQUS code

12 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I are relevant to the first two phases, while the third phase is in any case simulated by the command “element removal”. The hole expansion can simply be simulated by imposing and removing an uniform radial displacement on the nodes at the hole edge or the mandrel is directly modeling and its drawing across the hole. The presence of a lubricated split sleeve can be accounted for o not. Anyway the residual tangential stress evaluated only by the simple expansion FEM model agrees fairly well with the experimental values obtained by the Sachs’ drilling method. On the other hand the more complex models put in evidence some side effects of the process no present in the simplest ones. The introduction of the lubricated split sleeve, that mediates the contact between hole and mandrel, causes a loss of axial- symmetry: the maximum circumferential stress is reached in correspondence of the split, while the minimum one characterizes the diametrically opposite location. As a consequence, the fatigue resistance depends on the orientation of the split in the sleeve with respect to the load direction. As a consequence the best suggested orientation for the split is the one coincident with the load direction, conferring the same criticality to both sides of the hole. Besides the simulation of the actual mandrel drawing evidences that the hole is subjected to an no uniform radial expansion, but it is progressively crossed by the mandrel: the material flow in axial direction is appreciable and it causes a variation of the residual stress field along the plate thickness: the tangential component is minimum at the entry face, afterwards it increases up to the absolute maximum value at about the middle distance, keeping higher at the exit face. This suggests that a “double cold working” in reverse directions may be a solution for structures for which there is access from both sides. FEM analyses performed with a doubled mandrel drawing confirmed an improvement, in terms of compressive tangential component, of about the four fifth of the plate thickness. From the finite element analyses not improvement were expected for a double application of the process in the same direction. This can be interestingly too for applications with insufficient accessibility to both structure sides. Therefore focused fatigue and propagation tests were performed on “dog-bone” specimens in the aeronautical aluminum alloy 7075-T73, thickness 2.3 mm, with a central hole, simple and double cold worked. Fatigue tests highlighted the improved performance of cold worked holes and the extra enhancement effect of the process double application, both in concord and reverse directions. At higher number of cycles, the results show an increase of about 11% of the allowable fatigue stress for the case of conventional split sleeve cold working, while such an increment reaches about 18% and 25% for the case of double cold working performed in concord and opposite directions respectively. More interesting results come out from the propagation tests considering the two states of a initial defect lying on the entry (first drawing) side or on the exit side.

13 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I Experimental results, consistent with FEM analyses, put in evidence the strong retardation effect of cold working on crack propagation: higher is the compressive residual stress field, slower is the crack growth. In fact cold worked specimens life is, respectively, 10 and 30 times greater than life of the pristine specimens, respectively in the cases of initial defect at the entry and exit face. In the case of double cold working, life increases about 60 times for the entry face and more than 100 times for the exit face. Experimental results confirm also the positive effect of double cold working performed in concord direction in contrast, at least in part, with the FEM analyses. It is however conceptually difficult to find a justification for such a behavior. The evaluation of the residual stress field around a cold worked hole would have to be finalized to the fatigue life evaluation; in actual fact, only a few papers deal with this problem. In the present analysis the basic idea was the introduction of the residual stresses in a pre-existing fracture mechanics computer program. AFGROW was selected for this purposes, because it can account for the existence of residual stresses by calculating additive residual stress intensity factors Kres, which are added to the stress intensities (both maximum and minimum) caused by the applied loads. This will not change ΔK, but will change the stress ratio, R. The program is not able to evaluate the redistribution of the residual stress field due to crack propagation; this point, even if important, is not fundamental, because a large part of fatigue life is consumed in presence of small cracks, which would not cause significant residual stresses redistribution. Besides, redistribution is less important in the case of compressive residual stresses, because, contrary to the case of tensile residual stresses, compressive residual stresses can be present also in a cracked body. The experimental crack propagation data were so compared with those predicted by AFGROW: the predicted curve with no cold working shows a good agreement with the experimental data, while, a similar agreement can be obtained for the cold working case only by considering a small amount, about 30%, of the estimated residual stress field. AFGROW has a module, based on the "local strain approach", to calculate crack initiation life, so that the program can be used also for the evaluation of the total (initiation + propagation) fatigue life. In this case the predicted fatigue resistance of specimens with cold worked holes was unquestionably higher than the experimental evidence. Also In this case agreement was found by considering only 20% of the residual stresses (curve CW; AFGROW, RS 20%). In order to find possible justification for this behavior different additional tests were carried out at least to exclude some possible causes, such as a residual stress reduction (relaxation) proportional to the number of applied cycles. Five fatigue tests on cold worked holes were suspended respectively at 20, 40, 60, 80 and 100% of the mean fatigue life. But the results obtained with the Sachs’ method do not show a significant modification of the residual stress field due to fatigue loading.

14 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I Attention was so focused on potential improvements of the FEM model by means of the ANSYS code. A difficulty in cold working process modeling was certainly connected with the elastic-plastic behavior of the base material. Correct data are strictly necessary because of the highly positive-negative strains recorded in the material surrounding the hole during the cold working process. During cold working expansion a high circumferential deformation is imposed to the material near the hole, function of the adopted level of interference. When the mandrel leaves the hole the elastic spring-back of the surrounding zone usually causes also a re-yielding in compression of the hole edge material. As a first step toward the problem solution an experimental evaluation of the complete tensile-compressive stress-strain curve of the material, aluminum alloy 7075- T73, was obtained. The basic behavior under only tension or compression was compared with the results of more detailed tests. Some specimens were loaded in tension until a fixed value of total deformation (i.e. about 2, 3 and 4%) and, successively, they were loaded in compression. These stress-strain histories imposed to the specimens were selected in order to replicate the stress strain histories at different points near the cold worked hole edge. As a matter of fact, the yielding in tension occurs quite abruptly, while the compression branch exhibits a significantly more gradual behavior. The remarkable decrease of the yielding stress in compression following a yielding in tension attests a strong “Bauschinger’s effect” of the material, i.e. a strong diminishing of the compressive yield stress and a smoothing of the negative plastic part of curve. A direct comparison with the curves relevant to the most common models for the material plasticity, i.e. the multi-linear kinematic and isotropic hardening models, showed that the present material agrees fairly well with the kinematic hardening model used in the previous analyses , but remarkable differences are present in the negative part of the curve. In order to obtain an improvement respect to previous works higher accuracy on material model definition seemed to be strictly necessary to correctly evaluate the residual stress field around the hole: the first step was to optimize the material model with the base material tests as a reference. To define the inelastic behavior a non linear kinematic hardening model was selected, the Chaboche model. The ANSYS material option (CHABOCHE) typically uses the Chaboche’s model for simulating the cyclic behavior of materials. In fact is possible to superpose up to five kinematic hardening models and an isotropic hardening model to simulate the monotonic hardening and the Bauschinger’s effect and also complicated cyclic plastic material behavior, such as cyclic hardening or softening, and ratcheting or shakedown. In order to completely define a Chaboche numerical model 11 material constants must be defined. The challenge was to find correct values of these constants in order to obtain a good agreement with experimental tests. That represented an hard manual

15 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I task: consequently the optimization of the Chaboche material model was obtained by an integrated multi-objective and multi-disciplinary design optimization software platform, modeFRONTIER 3.1.3 code. In order to optimized also the process simulation a complete FEM model was developed: the employed high strength steel mandrel was directly modeled and it was drawn across the hole. The lubricated stainless steel split sleeve was interposed between the mandrel and the hole as in the real treatment of the hole. The implemented plate material model was the modeFRONTIER optimized Chaboche for the 7075-T73 aluminum alloy. The final reaming, tuned on the base of previous simpler analyses, was obtained by the ANSYS command “EKILL” (element kill), in order to correctly remove the inner layer of material. The simulation of the actual mandrel drawing highlighted useful information about stress and deformations during the different steps of the process: when the mandrel starts to cross the sleeve the gradual opening of the split follow correctly the movement of the maximum diameter of the mandrel and during the plate’s expansion it’s possible to record the evolution of the radial and tangential stresses in the material surrounding the hole edge step by step along the plate thickness. The three dimensional nature of the process introduce a loss of axial-symmetry due to the split sleeve and different stress field in the plate: also after the final reaming, the tangential component is minimum at the entry face, afterwards it increases up to the absolute maximum value at almost half of the plate thickness and slightly decreases at the exit face. The optimized Chaboche plastic material model and the more time consuming analysis, allowed a more correct quantitative assessment of the compressive residual stress field, that is reduced of the 24% on respect to the kinematic material model. Numerical results reliably matched propagation tests: what is more, observing the fracture surfaces of broken specimens, the double P shape of the crack fronts seems to trail the negative residual stress field along the thickness. As comparison some test were performed on the same specimen configuration, but with coupons processed by the StressWave method at StressWave Inc. Detailed process parameters were not divulgate: the only witness of this cold working phase was the small dimple on each side of the work piece. The second process step, hole drilling and reaming, was realized at the Department: the entire dimple and the surface upset were completely removed. Fatigue tests demonstrated an effect comparable with double split sleeve cold working. But when crack propagation tests were carried out on StressWave cold worked specimens at the same stress levels of split sleeve cold worked specimens cracks nucleated but do not propagated: fatigue cracks stopped when crack length was about 1 mm. Crack propagated up to the final failure of StressWave cold worked specimens only when the maximum fatigue stress was increased of the 13%. Results elaborated in the

16 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I form crack propagation rate as a function of stress intensity factor range, conservatively assuming a through crack front shape, indicated a strong reduction of crack propagation rate in Stresswave holes with a beneficial effect of present in a larger area respect to Split Sleeve cold worked holes. FEM analyses performed with a highly simplified geometry of the Stresswave indenters showed a good agreement in the residual dimple depth and geometry. Compared with the split sleeve cold working, a slightly higher compressive residual stress is obtained and a significant reverse yielding can be observed after the drilling of the hole. The region around the hole with beneficial compressive residual stress is approximately of the same dimensions than in the split sleeve cold working, but, as expected, the stress field is constant along plate thickness, as can be demonstrated by regular crack front shapes. The next goal will be the introduction of the obtained correct residual stress field in provisional software to correctly evaluate fatigue life of cold worked open hole specimens. When robust models will be achieved, others possible parameters variables would be considered in order to consecutively extend the results in more realistic joint structures. A consistent part of the research activity performed on this subject is described into the published papers [6], [9] and [10], fully reported in Section 4.1.2. Complete results are presented in the restricted reports [7] and [8].

Conclusions

Nowadays the use of cold working processes is usually limited to the solution of fatigue problems encountered in service or to introduce an imprecise beneficial higher security factor in structure critical points. To widespread these techniques is crucial a process deeper knowledge that allow to introduce it as a design parameter. At present time a new cold working process, the StressWave, seems to be competitive with the Split Sleeve Cold Expansion system for structures with accessibility from both sides, especially considering cutback in productive costs connected with the reduced number of process phases.

17 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I 2.3.4 EFFECTS OF “FORCEMATE” IN A LUG-FORK JOINT: WING TO FUSELAGE CONNECTION OF AERMACCHI M-346

Introduction

The experience acquired, both from an experimental and numerical points of view, in analyzing techniques able to enhance fatigue life in metallic structures suggested to investigate a practical application of another FTI process: the ForceMate®. The ForceMate technique promised to be a valid alternative to shrink or press fit bushing installation methods, with a significant life improvement. In fact the bushing is initially placed in the hole as clearance fit, and then using an oversized tapered mandrel, the lubricated ForceMate bushing is cold expanded into the hole. This process seemed to be really interesting because it could combine the two phenomenon previously analyzed: actually this high interference bush installation method is declared to create also a slight cold working of the main hole.

A new trainer aircraft, M-346, is currently under construction at AerMacchi (Venegono Superiore, Italy); all critical elements of this aircraft, i.e. those that would cause loss of an aircraft in case of non-detected damage, are designed on the basis of Damage Tolerance criteria, following the guidance of U.S. Joint Services Structural Guidelines, JSSG-2006. In the new trainer aircraft AerMacchi M-346 the connection between the main fuselage frame and the wing spar, that is certainly one of the most critical elements, is realized by a lug-fork joint. In the prototype configuration of the aircraft, each half-wing has three spars, connected to three main fuselage frames: both spars and frames are constructed by High Speed Machining starting from plates, thickness 42 mm, in 7050- T7451 aluminum alloy. The connection between each spar and the corresponding frame is made by two ‘lug-fork’ joints; the severe fatigue cycle spectrum to which these components are subjected obviously required a good design and a damage tolerance approach: The special ForceMate® bushings, produced by Fatigue Technology Inc., Seattle, are consequently interposed between the steel pins and the aluminum structure, in order to improve the component fatigue life. Also if applied to a unquestionably different geometry, this solution represented an interesting benchmark to evaluate the synergic effect of the interference fit and cold working techniques.

18 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I Activity

A large research activity was carried out in the framework of a cooperation between AerMacchi and the Department of Aerospace Engineering at University of Pisa, to support Damage Tolerance certification of the new trainer aircraft. Activity started with the characterization, from the Damage Tolerance point of view, of the material used in frame and spar construction of prototype aircrafts. Crack propagation and fracture toughness tests carried out on CT specimens put in evidence fatigue crack growth rate properties of the aluminum alloy 7050-T7451 fully comparable with those contained in the new database of NASGRO 4, the software selected by AerMacchi for the Damage Tolerance design of the new M346 trainer aircraft. Additional tests were carried out applying a load sequence representative of the load acting in the most critical lower wing lug. This was a necessary database to calibrate the Modified Generalized Willenborg (MGW) model that, among different options present in NASGRO, was selected to evaluate fatigue crack propagation of the M-346 structure. Subsequently, a larger ForceMate specimen, with the same dimension as the most critical lower lug in the spar-frame connection, was designed and fatigue tested under constant and variable amplitude loading sequences, representative of the aircraft current usage. An ABAQUS finite element simulation was conducted for the entire process in order to evaluate the effectiveness of the process itself: both the mandrel and the bushing were included in the model. Stress–strain curves of aluminum alloy 7050-T7451 for the specimen and steel 17-4 PH for the bushing were introduced in the model, while the mandrel was made of steel with only elastic behavior. The ForceMate was realistically simulated by drawing the mandrel across the hole, from the entry face to the exit face of the specimen, by imposing to it a translation displacement. Obtained strains were in good agreement with strain gauges measurements achieved during bush installation. Both the experimental and FEM results indicated that during bushing installation the hole is cold worked, with a additional beneficial effect from the fatigue point of view, besides interference fit mounting. To put into evidence single factor contributions, three comparative crack propagation tests on lug specimens were carried out under constant amplitude loading: a ForceMate bushing specimen, a specimen with a shrink fit bushing installed with very small interference (0.02 mm) after having removed a ForceMate bushing and one specimen without bushing. Results obtained show that, for the test conditions examined, ForceMate bushing increased crack propagation life about ten times respect an untreated and unprotected hole; but results relevant to the threaded specimen with only a reinstalled shrink fit bushing were more similar to results relevant to the specimen with ForceMate bushing.

19 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I This result highlighted that, in ForceMate bushing installation, the hole cold working effect prevails over the bushing interference effect; indeed the beneficial effect persists even if bushing is subsequently removed. Different fatigue tests were so carried out on lug specimens applying the design fatigue sequence and a 10% amplified one and with various initial defects introduced by spark erosion to check possible accidental damages. All specimens were instrumented by crack propagation gauges in order to monitoring front crack length during the fatigue tests. The experimental data were compared with the results predicted by NASGRO and by AFGROW. This software contains a recent and accurate solution for the Stress Intensity Factor of a corner crack in a pin-loaded hole and, as seen before, allows the introduction of residual stress effect. Results obtained show that actually cracks propagate very slowly if compared with predictions, particularly in the thickness direction; AFGROW predicts better the surface dimension of the crack, but in any case the result is very conservative. Observing lug fracture surfaces was evident an higher crack propagation rate in the lug surface direction respect to the lug thickness direction, contrary to what has been observed in numerous papers from the literature. To explain this behavior different hypotheses were formulated and checked by ABAQUS Finite Element analyses: results obtained with the Crack Opening Displacement method for parameters as pin elasticity and consequent bending with redistribution of contact pressure, friction coefficient in pin/hole contact and clearance in pin/hole contact, did not give rise to correction factors for the Afgrow SIF solution that can justify the behavior observed. This aspect, that could be also supposed to be in relationship with an asymmetric residual stress field produced by ForceMate bushings or the variable amplitude loading, was observed also in the specimen tested without any bushing under constant amplitude loading, so that such reasons must be excluded. Deducing finally that such behavior was a consequence of material anisotropy, additional crack propagation tests were carried out on square section specimens, loaded in the rolling direction; these specimens were machined by using broken lug specimens, so as to examine exactly the same material. At the optical microscope the material presented an highly irregular fracture surface; SEM analyses of the base material and the fracture surface highlighted a ‘pancake’ microstructure with normal size grains surrounded by very small grains. This microstructure, considered normal in the production of these materials, promoted nucleation of cracks in the lug longitudinal direction and slowed down the crack propagation rate in the short transverse direction. Additional tests were performed in smaller specimen extracted from lug specimen to investigate different crack propagations rate in the orthogonal directions. A significant

20 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I increase of the crack propagation rate in the short transverse direction was obtained for SIF higher than 13 MPa*m0.5, so in the range of lug spectrum stresses. The verified strong material anisotropy made consequently really difficult to correctly estimated fatigue life of this component by commercial fracture mechanic software for Damage Tolerance analysis. The main results obtained during this activity are described in the published papers [13] and [15], fully reported in Section 4.1.2. All the activity is described in the final report [14].

Conclusions

A large activity was conducted to support the certification of a primary structure of the M-346, the lug-fork wing to fuselage connection, treated with the ForceMate method. Despite complications encountered, the research unquestionably highlighted the great potentialities of the ForceMate process. Compared to shrink fit bushings ForceMate bushings increased fatigue life of the lug specimen more than five times and practically no fretting was observed. Focused tests and FEM analyses clearly demonstrated that, from the fatigue point of view, the cold working effect in a ForceMate bushing installation prevails over the interference fit effect. From a practical point of view, a consequence of the results obtained is that the AerMacchi M-346 lug-fork connection was over designed to support service loading. Introducing the Force Mate process the fatigue life of pristine specimens exceeded 10 times the design life. Even with the introduction of an initial flaw from 1.27 to 4 mm, the fatigue life was more than five times the design life. Actually is possible to reduce the wing spars from three to two, after a complete revision of the aircraft design.

21 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES I 2.4 CONCLUSIONS

A large research activity was conducted relevant to the fatigue resistance of structural components with holes expanded by means of the Interference Fit, Split Sleeve Cold ExpansionTM, StressWave® and ForceMate® techniques. Several FEM analyses, described in the presented papers, were performed to simulate process phases and effects, with the final aim of providing useful instruments for the estimation of the service life enhancement in high stressed aeronautical components. Obviously experimental tuning of FEM model is essential, specially for correct material model definitions. Simplest models demonstrated to be really effective, but more time consuming models can highlight important side effects and clarify unexpected experimental results. Good correlations between numerical and experimental strain gauge strains were always obtained; on the other hand a correct evaluation of stress fields around treated holes was not a trouble-free task. Qualitative evaluations of the high enhancement effects on fatigue life of the analyzed processes were experimentally verified. However crack growth predictions models implemented into commercial codes need an extra improvement to be considered reliable instruments in Damage Tolerance analyses of critical structures.

22 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I 3. INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES

3.1 INTRODUCTION

The need to design new commercial aircrafts to replace the existing ones at the end of their design life and to increase the world aircraft fleet that will grow in the next twenty years, has forced the aerospace International Community to investigate the possibility of obtaining more efficient structures in the cost and/or weight reduction field, with the same or better safety levels than the actual ones. Introducing new aircraft structures, cost and weight savings are pursued by an integrated design. The aerospace industry uses riveting as a fastening technology to join the aircraft panels of fuselage, wings and vertical and horizontal stabilizers. Nowadays great efforts are made to replace this standard joining methodology by new technologies which would provide considerable weight saving and cost reductions incorporating integral structures in aircraft constructions. But any new material or new joint technology for airplane fuselages in areas where damage tolerance is required, can only be accepted if the crack growth behavior is not worse than the standard 2024-T3 material and standard riveting technology. In this framework the Department of Aerospace Engineering started a cooperation in the national research program COFIN 2002 for the “New Technologies for Pressurized Fuselages in Medium and Large Size Aircraft”; the project final aim was to investigate the possibility of obtaining more efficient and reliable structures introducing innovative technologies or materials. During development of this research activity major attention has been paid to deeply investigate new welding technologies, on which is currently focused the interest of both aeronautical and aerospace international industries. Among others the Friction Stir Welding (FSW), a new process developed by The Welding Institute in the early nineties, is proved to be a very viable joining process also for all the aluminum alloys difficult to weld by conventional fusion procedures. Sheets and plates with a maximum thickness of 100 mm of 2XXX, 5XXX, 6XXX and 7XXX series aluminum alloys has been welded, demonstrating that it could be an effective process for replacing bolted and riveted joints. The structural component where attention has been mainly focused is the fuselage: this technology can be intensively employed in joining e.g. stringers or clips to the skin sheet or in butt joints of extruded panels with integral stringers with a considerable reduction in weight and manufacturing. But if Friction Stir Welding is considered for primary aircraft structures different parameters must be analyzed and specially the in-service behavior of these systems have to be clearly defined. The final aim will be to develop an accurate and efficient

23 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I crack growth life prediction methodology for this specific joint type that, representing the basis for scheduling periodic inspections and for maintenance programs, will ensure flight safety at a lower cost.

3.2 THE “FRICTION STIR WELDING” PROCESS

Friction Stir Welding (FSW) is a solid phase joining technique developed by The Welding Institute (Cambridge, UK) and patented by TWI in 1991. Because of its solid phase, FSW allows the high quality joining of materials that have been traditionally troublesome to weld conventionally such as high-strength aerospace aluminum alloys like 2024 or 7475. The present status of FSW allows one to weld all aluminum alloys in thicknesses ranging from 1.2 to 100 mm, and even dissimilar aluminum alloys can be joined. A friction stir butt weld is produced by plunging a cylindrical rotating, non- consumable tool into the faying surfaces of the two plates to be joined (the work). The parts have to be clamped onto a backing bar in a manner that prevents the abutting joint faces from being forced apart. A generic tool consists of a shoulder and a profiled pin emerging from it. To initiate the welding process, the rotating tool is slowly plunged into the faying surface until some substantial portion of the shoulder is in intimate contact with the surface of the work-piece to avoid expelling softened material. In order to make a full penetration weld, the tool must be designed so that the bottom of the pin is near, but not at, the bottom surface of the work when the shoulder makes contact with the top surface. Hence, welding of various thickness work-pieces requires tools with different pin lengths or a special tool with a pin element of adaptable length or continuously retractable. Once the plunge is completed, traveling of the rotating tool along the joint line results in the formation of the weld. In fact, by the friction and the pressure due to the relative motion, the metal edges warm and become plastic without reaching melting point. So the softened weld metal is extruded around the pin and then forged by the trailing edge of the shoulder. The process can be visualized as an in-situ extrusion. The extrusion and forging operations, coupled with tool rotation, are responsible for the characteristic microstructure that results from the FSW process. The weld microstructure features can be separated into two broad categories: the thermo- mechanically affected zone (TMAZ), and the heat affected zone (HAZ). The TMAZ might be considered analogous to the fusion zone of a conventional weld except that, instead of being melted, the material in the TMAZ has been plastically deformed by the friction stir welding tool. The centre of the TMAZ, the “nugget”, has a dimension similar to the diameter of the pin and is characterized by a fine, relatively equiaxed and re-crystallized grain structure.

24 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I The HAZ which will clearly lie closer to the weld bead, is similar to HAZs resulting from conventional, fusion welding processes: the material has experienced a thermal cycle which has modified the microstructure and/or the mechanical properties. However, there is no plastic deformation occurring in this area. It should be noted that the FSW process is not symmetric about the centerline: the tool rotational velocity has a concord or opposite directions respect to its traveling direction at the two tool sides, defined respectively as advancing and retreating side. During the last years extremely attractive results have been obtained when FSW has been used on high strength aluminum alloys. This brought “Friction Stir Welding” to be recently identified by leading aircraft manufacturers as “key technology” for fuselage and wing manufacturing. As a matter of fact a lot of process variables can affect the quality of friction stir welds: tool design, tool rotational speed, travel speed, tool heel plunge depth, relative position tolerances, etc. and the base material as well. Although the vast majority of friction stir welds are expected to be flaw-free, it is not always possible to assume this. The principal aims are to fix welding procedures and material selection in order to enhance the confidence in the design and application of friction stir welded joints. The Department of Aerospace Engineering at the University of Pisa developed a research activity on this argument to study in a deep and methodical way some fundamental aspects, in conjunction with the international industry aims.

3.3 “FRICTION STIR WELDING” FOR AERONAUTICAL APPLICATIONS

3.3.1 FSW ON 6013 ALUMINIUM ALLOY: IMPACT OF A DIRECT WATER COOLING SYSTEM

Introduction

An important study was performed in the context of a cooperation between the Department of Aerospace Engineering at the University of Pisa and EADS Corporate Research Centre Germany of Ottobrun, Munich: this work was included in a wider project, WAFS, focused on the evaluation of performance of Friction Stir Welded joints and the optimization of process parameters. As introduced before the FSW technology can be intensively employed in manufacturing fuselage welded integral structures: the present activity branch mainly evaluated the welding process in butt joints of extruded panels with integral stringers on

25 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I the aeronautical aluminum alloy 6013-T4 and the possible introduction of an innovative fluid-gas cooling device to reduce residual stress into the joint. Attention was specially focused on the evaluation of the static and fatigue properties of the friction stir welded joints, of the residual stress and on micrographic analyses of the joints. Additional tests were performed to critically evaluate the effect of the possible presence of defect into the welded joints. Above all this study represented the first step toward a deeper process knowledge, including process typical parameters, welding distinctive features, characteristic defects, significance of mechanical tests and inspection methods to critically evaluate this new technologies.

Activity

The present investigation was performed using 3 mm thick Al-Mg-Si-Cu 6013 aluminum alloy, in the naturally aged condition T4. This material is often presented as a possible alternative to the widely used AA2024-T3 Alclad in a corrosive environment. The main advantages of this alloy are a lower density a reduced fatigue crack growth rate, a higher yield stress and higher fracture toughness; besides the higher elongation to failure often allows to obtain a formed sheet by only one mould. This material was consequently selected to realized extruded panels 3 mm thick with integral stringers; in correspondence of the welded joint line, obviously parallel to the plate extrusion direction, the local thickness of the panels was increased to 4 mm in order to furnish sufficient material. An initial large amount of testing and parameter tuning is necessary to create acceptable Friction Stir Welded joints. In particular it is important to chose tool rotation and welding speed in order to minimize heat input which might introduce weakening effect in the HAZ. In the present activity welding parameters had been selected by EADS Research Germany with an innovative designed and manufactured combined fluid-gas cooling device. The water cooling device works on the weld bead and the gas nozzle “cleans” the material in front of the weld seam in order to prevent excessive water at the joint line interface. The purpose of this system was to minimize residual stress effects like distortions, possibly increase the corrosion resistance of welded joints and improve the overall strength properties of Friction Stir welded materials (smaller heat affected zone and enhanced artificial aging response). Two complete welded panels were obtained under some slightly different conditions, giving rise to two different series of tests. The first one was welded under the nominal condition parameters; the second panel was intentionally welded with a

26 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I misalignment between the tool axis and the abutting line, in order to simulate a possible error in tool positioning in a large aeronautical structure, as the fuselage will be. The plates welded were subsequently sectioned transverse to the welding direction to obtain two series specimens. Global mechanical analyses highlighted high performance of Friction Stir Welded joints with a maximum failure stress of about 75% of the base material final stress; this value is much more impressive if compared with the typical failure strength of joints realized with classical fusion welding technologies. Interesting observations were obtained during elongation tests. Standard measurement of the percentage elongation hypothesizes uniform deformations along the entire gauge length: thus a very low value, 3.8%, of elongation to failure was obtained in FSW joints. But, at the joint final failure, the base material is still in the elastic field. So the apparently brittle fracture exhibited by the welded specimen was only a result of the presence of a plastic strains limited to the weld bead, in which instead a local peak value over 16% was obtained. Fatigue tests evidenced high performance of the 6013-T4 FSW joints, with a resistance of about 70-80% respect to the parent material. This is a promising result from a design point of view, instead of the factor two or more in fatigue life obtained with traditional welding technologies. Results could be separately processed for the two series in order to show the effect of pin and butting line misalignment. But surprisingly results were not so different: instead tests relative to the first series, with nominal parameters, showed a slightly higher scatter and a lower fatigue limit in the range of low stresses. On the contrary each series showed a characteristic fracture surface feature, clearly correlated with its fatigue performance. So the activity was mainly concentrated on analyzing the welding quality and the weld features that could provide very useful information about this new process. In the so-called “solid state” FSW process, typical welding defects, like porosity, shrinkage cracks and segregation, practically disappear. But, although the FSW process showed to be very robust in terms of the range of welding parameters that will produce acceptable welds, the weld bead can contain different types of defects such as discontinuities, micro voids and micro-structural imperfections. At first, metallographic sections were used to try to detect flaws. The photos showed the characteristic joint microstructure with the banded structure of the nugget which forms the so-called “onion rings”: these are traces of the softened weld metal extruded around the threatened pin; in any case, for the magnification employed, no voids, joint line remnants and root flaws appeared. Fracture surface showed in both series some characteristic surface marks in the shape of chevrons. The chevron mark direction is always opposite to the welding direction, with a pitch between two successive signs equal to the feed per revolution.

27 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I Analyzing instead at the optical microscope the lateral surfaces of the specimens of the first series, in some case a visible void was clearly identified. This “tunneling” defect was intermittent but clearly visible also on the fracture surfaces on the retreating side of the weld, showing a clear delineation between the region of the nugget and the remainder of the TMAZ. Classic porosity in fusion welds is caused by gas bubbles, or by shrinkage at the time of solidification, and clearly neither of these processes is possible in FSW. Such defects should therefore be referred to as “volumetric flaws” or “voids”. In fact, welding at high traveling speeds, the material receives less work per unit of weld length, i.e. fewer tool rotations per unit length. Under such conditions, the plasticized material may be cooler, and less easily forged by the shoulder, resulting in voids remaining unconsolidated. So it can be deduced that the cooled FSW process still needs a proper parameter arrangement, as an higher tool revolution with the selected travel speed to overcome the evident heat losses. The second series panel, obtained with the same parameters set up, shows instead a different failure mode. The introduced pin misalignment had been expected to give rise to some root defects, even if the external visual examination failed to detect the presence of root flaws. The visual and microscopic examination of the fracture surface revealed a line, slightly brighter in color than the parent material, running the entire width of some specimens. The line was approximately 0.1 - 0.13 mm wide and positioned at the weld root. It does not necessarily represent an absence of any bond, indeed some regions of intermittent or weak bonding may be present. This type of defect represents the most common flaw observed in a single-sided Friction Stir weld. This type of defect is really difficult to be identified using standard NDI methods because of its nature, which had been shown to be one of intimate contact with not fully or poor strength bonding. This defect appears at the specimen surface only when it is subjected to a progressive localized overstressing. So bending tests for ductility of welds were conducted, according to the normative ASTM E 190, with the weld root positioned in the tensile side. When specimens reached a bending angle of about 90° degrees, clearly visible flaws appeared on the root side. Since the ASTM prescribes a flaw-free 180° degrees bending angle, the weld had to be defined as flawed. In the Damage Tolerance analysis of integral structures is basic to define the dimension of the detectable defect. At the present stage, although the Friction Stir welds seemed to be tolerant to the presence of root flaws without a significant performance penalty, only periodic destructive bend tests appear efficient on monitoring ongoing process quality. Obviously the selected “cooled” FSW process still needs a proper parameter arrangement before starting production.

28 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I The integrity of mechanical structures can be strongly influenced also by the presence of residual stress distribution particularly when they experience in service a significant fatigue loading. Friction Stir Welding is seen as a promising method because its limited heat input could highly reduce joint residual stress distributions, specially if compared with standard fusion welding techniques. However, it must be consider that aluminum alloys have a very high thermal conductivity and coefficient of thermal expansion and these aspects, combined with the relative low yield strength and a rigid clamping, can cause a certain amount of residual stress during the cooling of the welded area after FSW. One of the main goals of the cooling device employed was to strongly reduce the amount of geometrical defects and residual stresses in a welded joint. Actually the visual analysis the welded panels showed a very limited amount of geometrical defects if compared with others no cooled FSWelded or laser welded panels examined in other works. Residual stresses were measured by using a destructive method: a panel was instrumented with strain gauges positioned in a section transverse to the weld bead on the tool side. Subsequently the panel was sectioned with a milling machine transverse to the welding direction in order to allow the relaxation of internal stresses. At this point the principal aim was only a general evaluation of thermal residual stresses specially compared with those that compete to others welding technologies. In the case of the 6013-T4 aluminum alloy welded with this particular cooling device, a maximum value of 96 MPa was measured, slightly out of the weld zone, on the retreating side. This value is fully comparable with a value obtained in a classic FSW process. So the cooling process did not appear so effective from the residual stress point of view even if the residual strains observed were very low.

Conclusions

Different analyses carried out on 6013-T4 Friction Stir Welded joints realized on integral extruded panels by EADS Germany represented a first chance to checking process knowledge. Good static and fatigue performance increased process confidence also if two typical defects were identified. Different failure mode were analyzed and clearly connected with an insufficient heat input. Compared with others FSW or laser welded panels previously examined, present ones showed a limited amount of geometrical defects thanks to the properly designed cooling device but the process still require an optimization of welding parameter in order to definitely achieve no root weld defect due to bad or incomplete bonding.

29 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I 3.3.2 FSW ON 2024 ALUMINUM ALLOY: EFFECT OF FLIGHT LOAD SPECTRUM

Introduction

Loading history experienced by aircraft structures is usually a variable amplitude loading: Different types of load sequences induce a number of different load interaction effects, retardation and acceleration, which can result in significant variation on the crack growth rate. Unfortunately, there are so many interaction effects between load cycles of different amplitudes that, even if some of them are quite well understood (at least from a phenomenological point of view), carrying out accurate quantitative predictions is still a formidable task. One of the key issues during the material selections for light weight aeronautical structures is the fatigue crack propagation behavior of cracks under a defined in-service environment. The medium strength structural 2024-T3 aluminum alloy is the typical candidate for the fuselage skin, currently in use due to his attractive combination of mechanical strength and particularly good damage tolerance. As seen before the Friction Stir Welding can be efficiently applied to this structural aluminum alloy, but before introducing this new joint technology on primary aircraft structures is indispensable to verify his fatigue crack growth behavior under flight loading conditions and to elaborate prediction methodologies for this specific joint type. Currently a very limited amount of data on fatigue crack propagation in friction stir welded joints exists. Most of the work on fatigue of aluminum friction stir welds has been restricted to the generation of fatigue life S-N data (stress levels vs. cycles to failure curves). What is more there is a complete lack of information regarding the fatigue crack propagation in Friction Stir Welds under variable amplitude loading and flight loading conditions, both from an experimental as well as from a prediction point of view. In the present investigation an exploratory test program was carried out at the Institute for Material Research, Mechanics of Materials and Joining Technology of the DLR (German Aerospace Centre), in Cologne, Germany, to improve the general understanding of the loading sequence effects in Friction Stir Welded joints. The first aim was to offer a basic experimental data base to be used as a first benchmark of the existing crack propagation prediction models. The final objective was to elaborate an easy, practical and reliable engineering approach to perform crack growth life predictions of Friction Stir Welded structures under in-service loading. However experimental residual stress measurements on

30 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I Friction Stir Welded panels will allow a more detailed application of fracture mechanics principles in crack growth evaluation.

Activity

The base material used on the entire experimental program is the widely used aluminum alloy 2024-T3, 4 mm thick. All the Friction Stir Welds were performed on a conventional milling machine at DLR as single side butt joints parallel to the rolling direction of the sheets. Some preliminary tensile tests performed on the base and in the Friction Stir Welded materials shown that the DLR optimized welding parameters give rise to strongly reliable joints. In order to obtain a sufficient crack propagation database 10 CCT base material specimen were used as reference and 15 specimen were obtained from FSW panels with the bead in the centre of the section, transverse to the loading direction. The initial defect in the FSW coupons was positioned 5 mm out of the weld centre on the retreating side, which was previously demonstrated to be the most critical part of the joint. All the fatigue crack growth tests were performed according to ASTM E 647-00 for CCT specimens. Some preliminary tests, both in base and Friction Stir Welded material, were conducted at nominal applied stress ratios of R=0,1 and R=0,8. Constant amplitude test are necessary for tooling prediction model on the basis of the actual base material and for put in evidence the effect of the residual stress present on the welded material. The maximum stress ratio R was selected in order to be sure that no closure will be present. Subsequent coupons were tested under simple variable amplitude load sequences such as single overloads, overloads and under-loads, overloads and compressive under-loads, and under standard flight load histories, FALSTAFF and TWIST. FALSTAFF and TWIST are quite complex load sequences pertaining respectively to fighter aircraft wing bending and to the wing root of transport aircraft: both were selected for collecting operative database on real loading conditions for welded primary structures. Different maximum loading levels for the flight spectrum were tested, between 120 and 245 MPa: FALSTAFF with a maximum of 196 MPa showed the more repeatable fatigue crack propagation life and therefore was set as maximum experimental load. Twist was set to a maximum of 120 MPa in order to avoid too fast crack propagation. The simplified sequences instead include individual blocks of a short length and always identical sequences of load cycles within the block in order to avoid complicated

31 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I superposition of short term and long term sequence effects as they occur in random loading. Every selected variable amplitude loading histories cause different significance sequence effects, moving from the strongest to the minimum retardation effect. In the spectrum with overloads and negative under-loads the maximum and minimum load levels were selected in order to have the same lowest load ratio present in the Falstaff spectrum, Rmin = -0,264. Constant amplitude tests were controlled by an Instron software while variable amplitude tests and standardized tests were controlled respectively by a WaveMaker 32 and a Fast Track software. For tests with a loading spectrum entirely in traction, during crack propagation a Potential Drop device was employed for crack monitoring, checked with periodical optical measurement of the crack tip. In partially compressive spectrum loading tests anti-buckling guides were indispensable, so a crack optical monitoring was applied. In order to achieve correct optical measurement of the crack length Falstaff tests were interrupted two times in 200 flights, flight 32 and 173, and Twist tests after the 5 maximum spectrum loads, flights 501, 1656, 2856, 2936, 3841. The complete program allowed a comparison between base and friction stir welded material for every loading spectrum. For each spectrum one base metal curve and two FSW curves were compared in order to obtain also information about scatter in welded structures. As general trend FSW specimens showed a faster fatigue crack growth rate as was expected by considering different material properties on the welded material and the presence of residual stresses in welded panels. Taking into account only mean values fatigue life of welded structures is reduced of about 40% respect to base material. Observing crack surfaces overload markings are clearly visible only for Falstaff spectrum loadings, both in base and welded materials. After a certain length all cracks tend to turn at 45° respect to the initial plane, both in base and welded specimens. Additionally in every friction stir welded specimen a tendency of cracks to turn in a plane immediately out of weld beads was observed. Crack propagation on a welded structure can be strongly influenced by the presence of residual stress, mainly when it experiences a significant in-service fatigue loading. Experimental residual stress measurements obviously become fundamental to obtain more detailed crack growth evaluation. In the present activity residual stresses measurement was performed at the Department of Aerospace Engineering at University of Pisa by means of a destructive method that had been demonstrated to be really effective. In this application strain

32 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I gauges were placed on the tool side in a section parallel to the weld bead, 5 mm out of the butt line in order to measure residual stresses in the nearest section to the crack propagation plane. The very rigid clamping devices employed in Friction Stir Welding exerts a high restraint on the deformation of the welded plates which impedes the contraction due to cooling of the weld nugget and heat affected zone in longitudinal as well as in transverse direction, introducing transverse and longitudinal residual stresses. During present activity most attention is focused on stresses that could affect crack propagation rate of crack disposed along the weld bead. Since longitudinal residual stresses were expected to be higher than the transversal ones, strain gauges were oriented both in welding and transversal directions. With this distribution, more correct results can be obtained combining strains measured in the two directions by means of biaxial elastic relationship. The cut parallel to the weld bead was performed starting from specimen centre moving in both directions to the edges of the coupon, like a fatigue crack on previous tests and cut strains were so recorded for different total cut lengths. Besides milled cut cannot be assumed as narrow as a crack front, present results can give a reasonably accurate value of the amount of residual stress acting on the welded structure. Although Sx maximum value, around 35 MPa, is not so high due to coupon finite dimensions, residual stress fields will be taken into account in crack growth prediction models. Calculations were performed with two fracture mechanics software, AFGROW and NASGRO 3.0. To keep the model simple and reliable some basic assumptions were introduced. The tested base material showed a slightly faster crack growth rate than the NASGRO database, but, for sake of simplicity, the 2024-T3 orientation T(L) material stored in NASGRO database was selected in subsequent predictions. As previously underlined, most evident effects acting on a structure subjected to variable loading histories are retardation and acceleration of fatigue crack growth, the so called “interaction effects”. In welded structures, fatigue crack propagation is also affected by the presence of residual stresses. Those two macroscopic effects will be firstly separately simulated and than combined. The implemented retardation models of these fracture mechanics software were firstly tested on the base material specimen with the easiest test program loading spectrums. The retardation model that gave rise to the best results was subsequently used to predict crack propagation experiments under more complex loading histories. In welded specimens, only AFGROW was used in prediction calculations due to his capability to take into account also residual stresses by calculating a residual stress intensity factor, Kres, that the code simply add it to the stress intensity factor due to the external load.

33 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I The so estimated lives for the base material were in good agreement with the actual lives showing the viability of the assumptions made in the analysis. Particularly for loading spectra with a predominant influence of overloads the Willenborg model with a shutoff overload ratio RSO = 3 gives the best results, while for loading spectra with a predominant influence of under-loads the best predictions were obtained without using any retardation. For FSW panels prediction were good but not-conservative for service loading spectra. For this applications the introduction of the welded material as reference database could probably improved the results, taking into account the changed mechanical properties and crack growth rate at different stress ratios. More details of the results obtained were published in [22], in German language, not reported here, and in [26], fully reported in Section 4.2.1.

Conclusions

A simple engineering approach which is based on a relatively solid background of fatigue test data for various test conditions was developed: it may provide a practical and reliable basis for the fatigue analysis of integral structures under in-service loading conditions, by using widespread fracture mechanics software. Fatigue crack propagation tests were performed on 2024 friction stir welded panels not only under constant amplitude loading but also under variable amplitude and standard flight loading histories. The two most important effects affecting the fatigue crack propagation under complex and in-service loading conditions of welded structures were identified and firstly separately simulated: the residual stresses and the retardation and acceleration effects. The single solutions of the two different problems were then superimposed to predict the fatigue crack propagation of friction stir welded specimens under more complex loading spectra and under flight simulation. The experimental measurement of the residual stresses and the inclusion of the retardation models of the software packages improved the quality of the predictions considerably.

34 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I 3.4 FRICTION STIR WELDING” FOR AEROSPACE APPLICATIONS

3.4.1 FSW ON “WELDALITE” ALLOY: THE FAST2 ALCATEL ALENIA SPACE ITALIA CRYOGENIC TANK

The present part of the activity concerns the results of a large research activity realized thanks to a partnership between the Department of Aerospace Engineering at University of Pisa and Alcatel Alenia Space Italia for the qualification of an innovative welding technology, the Friction Stir Welding by “Retractable Pin Tool”, on a new aluminum-lithium alloy for aerospace applications. Lightness and hermeticity associated with good mechanical properties are the main design requirements for an aerospace structure. In the continuous research on an high performance material for these applications, great attention is currently focused on the innovative 2195 Al-Li alloy, commercially known as WeldaliteTM. This alloy is mainly characterized by a reduced lithium percentage compared with previous Al-Li alloys; enhanced mechanical properties are so obtained and combined with a significant reduction in specific weight. WeldaliteTM is a weldable alloy and can also be efficiently used at cryogenic temperature. So it appears as an excellent substitute of others widely employed aluminum alloys, like 2219, on the construction of space tanks launch vehicles. Respect to the aeronautical industries, bigger efforts has been made on the improvement of higher quality welding technologies; in fact only welded joints can realize perfectly hermetic pressurized structures. Among others processes “Friction Stir Welding” has been currently selected by aerospace industry as the milestone technology for the next future space structures and is becoming a space industry reality. As a target of deeply innovation Alcatel Alenia Space Italia acquired a Pilot Plant allowing both hoop and longitudinal welding and started to optimized the welding process parameters on the promising 2195-T8 aluminum-lithium alloy. The final aim was obviously the qualification of the welding process, on the basis of Alenia quality standards, in order to use this process as a possible concurrent of TIG (Tungsten Inert Gas) and VPPA (Variable Polarity Plasma Arc) welding processes actually employed. One of the main problem faced during the activity was the presence of the hole at the end of the welded line due to the final retraction of the pin tool. On planar welded joints the extraction hole can be positioned on an extension of the welded plates and subsequently removed. But this “defect” cannot obviously be present on a circumferential welded joint of a pressure vessel.

35 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I This difficulty was overcome adopting an innovative FSW tool called “Retractable Pin Tool”. At the end of the weld bead the pin can be progressively retracted into the shoulder without leaving the classical “key hole”. The activity was successfully concluded by the design, manufacturing and testing of the 2195 Al-Li Cryogenic Tank Demonstrator integrally Friction Stir welded on the framework of the FAST 2 (Future Advanced Space Transportation Technologies) project, financed by the Italian Space Agency. To develop the program the Aerospace Department utilized the financial support of the COFIN 2003/2004 program.

Activity

The large experimental activity conducted had as final aim the qualification of the Friction Stir Welding process on the innovative Weldalite aluminum-lithium alloy. In order to obtain a complete characterization of the process several aspect must be analyzed both for the base and welded materials, starting from general mechanical properties, fatigue resistance, crack propagation fracture toughness, stress-corrosion properties, residual stress and non destructive testing capable to identified typical defect of this technology. The fist part of the activity was obviously focused on to correctly define base material properties. In fact aluminum lithium alloys has been investigate for more than 50 years but only during recent years had find important applications in the aerospace field. Lithium is known to increased in aluminum alloys the static strength, the elastic modulus and to diminished the specific weight, all properties really interesting in a minimum weight design. Actually their introduction on primary structures has been limited by some weaknesses on fatigue and stress corrosion resistance, particularly on the short transverse direction. These problem are declared to be overcome in the new generation aluminum-lithium that are weldable and maintain high properties also at cryogenic temperature. Mechanical properties evidenced high yield and static strength, totally unaffected by preliminary corrosive or stress-corrosive environments, but with a relatively low elongation to failure. Fatigue resistance was also really high but an unexpected behavior was highlighted. Fracture surface were really uncommon for aluminum alloy, with long secondary crack propagated also in the loading direction and with a wooden final fracture. As seen in the previous activity this like “pancake” structure makes really difficult to predict crack propagation, complicating structure design. Some fatigue tests were consequently monitored by a camera during fatigue life, and finally it was possible to declare this surprising behavior affects less than 5% of the fatigue life, strongly reducing provisional problems.

36 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I This aspect could however highlight an alloy weakness on the short-transverse direction. Small coupons were so designed to be extracted from the thicker available plate to test fatigue properties in this direction. Results, despite the “wooden” surface of broken specimens, didn’t demonstrated a significant resistance reduction. Static tests on the first friction stir welded specimen highlighted a yield and ultimate strength respectively of the 50 and 70% respect to corresponding values of the base material, unaffected by the environment as the base material. That means that in FSW the “joint coefficient” must be evaluated on the basis of the yield strength, being the most influenced factor. In fact the ultimate strength of the welded material is lower than the yield strength in the base material and, as observed in the previous analysis, that brings to plastic strains limited only to the weld bead. Elongation local measurements showed a value to failure double than the base material. Fatigue tests compare optimally with base material and are unquestionably higher respect to result obtained with the other consolidated welding technology, as TIG and VPPA. Fatigue test are not usually a design parameters for aerospace structures, but they evidenced weaker points where introduced artificial defect for leak before blast analyses. In fact all fatigue cracks appeared in the terminal zone of contact of the pin shoulder, on the border of the Thermo-Mechanically Affected Zone. This aspect was also confirmed by the micro-hardness tests that put in evidence a minimum of the Vickers hardness in this position, more marked on the retreating side. Crack propagation tests were consequently performed in the base material, as test reference, both in the LT and TL directions, at two different values of the stress ratio R. Results well compared with the 2024-T351 curve, milestone in Damage Tolerance analysis. Differences registered between the two orientations were small enough to considered only one mean curve of the Weldalite base material. What is more results can be interpolated with the Walker law, that will be very useful on crack growth prediction model. Also in this case the “wooden” aspect was clearly visible in the fracture surface, specially in the TL direction. In Friction Stir welded panels obviously different positions of an initial defect must be analyzed: propagations rate of defect parallel to the weld bead was slightly higher at the centre of the weld bead, instead practically the same value of the base material were obtained for defect positioned at the retreating and advancing sides, in correspondence of the end of shoulder contact. Defect propagated regularly without changing the initial plane. For each test condition two nominally identical tests were conducted using panels extracted from the initial and the final part of the weld line, in order to highlight possible mechanical property variation.

37 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I Globally it’s possible to affirm that the welding doesn’t affect base material properties for a defect positioned parallel to the joint line. Instead for defect positioned transversal to the weld bead a marked increase of the crack propagation rate was observed, also if the crack propagate toward parts less affected from the welding heat input. This effect is obviously connected with the residual stress field that increased crack propagation till the crack doesn’t reach dimensions able to relax the internal stresses. Residual stresses measurement was so performed with the destructive method already presented in different points along the weld line. A peak value of 130 MPa was obtained at the centre of the weld, that quickly decreased moving award from the centre line. The obtained value well compared with conventional welding technologies and positive stresses are present in a smaller area. The stress intensity factor connected with this residual stress field was calculated by the Tada-Paris equation, that take into account also the relaxation of the stress during crack propagation. By means of the base material database obtained previously and taking into account the effect of residual stresses so calculated, very good prediction of the crack propagation were performed. Using this consolidate results the activity focused on the effective welding process on longitudinal and circumferential joints to find possible weak points, first of all the closure hole. The new technology of the “Retractable Pin Tool” allows to obtained circumferential welding without leaving a marked defect on the joint line. In order to avoid problems connected with the possible different mechanical properties of the initial and final part of the weld line, the joint was obtained passing, after a first complete turn around the structure, a second time on the initial part and progressively retracting the pin into the tool shoulder. Some detailed investigation were obviously compulsory to eliminated doubts on the technology reliability. Metallographic section of the single and double pass weld bead didn’t evidenced significant changing on microstructure, except for a larger nugget area in the double pass welding. Fatigue tests with simple coupon centered on the different points of the joint line didn’t evidence a reduction of performance: instead results relative to the double pass welding had a reduced scattering of data. The tool second pass probably rework the material obtaining a more homogeneous microstructure. Specimens centered on the closure had a small reduction of fatigue resistance, but still in the scattering of the single pass data.

38 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I The real differences were present on the fracture surfaces. Both in single and double pass specimens fatigue cracks nucleated in correspondence of the final contact of the tool shoulder, with the same probability on the retreating and advancing sides. At the closure point instead cracks nucleated at the centre of the weld bead: the starting point could be clearly connected in each case with some process traces, in correspondence of the screwed pin profile. For aerospace applications the selected Non Destructive Inspection method is expected to clearly identify process defects. As seen in the previous analysis that is not an easy task. The Eddy Current method already employed in Alenia was so compared with a new ultrasound technique, the Multi Phased Array. In order to provide a reference database some panels were tested at different percentage of the fatigue life. When a defect was evidenced the panel was monitored at different steps of life, to evaluate the evolution of the defect and the accuracy of the inspection method. Eddy currents demonstrated to be effective in find superficial defect, but Multi Phased Array, although still not optimized, were capable to find out internal defects giving as output a clear image of the section view. All the performed activity allow the optimization and the certification of the welding process. Finally, as technology demonstrator, was successfully realized the cryogenic tank FAST2: the cylindrical part was obtained joining with conventional Friction Stir Welding four pocketed panels curved with Brake Forming; the two hemispherical domes, obtained by Spin Forming, where consequently connected with the Retractable Pin Tool FSW with two circumferential Friction Stir Welded joints. The main features of this new application and the results of the tests required for the process optimization and qualification are described in the presented papers [16], [18] and [21], fully reported in Section 4.2.2, and in the internal reports [20] and [25].

Conclusions

The large research activity performed in collaboration with Alcatel Alenia Space evidenced again the high potential of the Friction Stir Welding Process coupled with the use of advanced alloys like 2195 Al-Li alloy WeldaliteTM. The joints didn’t present internal defects, as confirmed by the optimal mechanical properties that well compared respect to more traditional techniques already in use in Alenia Space. Fatigue tests didn’t find significant weaker points, and defect nucleation was clearly identified by the selected Non Destructive Inspection methods.

39 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I The process is definitely very promising and can allow further safe optimization of the structural mass in space structures with all the related benefits in terms of costs and increased launchable payload capabilities. The effective solution was so employed for the production of the FAST2 cryogenic tank. As a consequence of the good result obtained the Friction Stir Welding Process has been introduced in the preliminary design of the CEV (Cabin Exploration Vehicle), the pressurized space vehicle for human mission on the Space Station with possible extension to inhabited mission on the Moon and Mars.

40 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES I 3.5 CONCLUSIONS

During previous years aeronautical industries ostracize the introduction of fusion welding as joint technology, but nowadays the new “solid” Friction Stir Welding process represent a promising candidate thanks to his limited heat input. Although many industries started to look at the process with increasing interest, enough information on the mechanical properties were not available to allow immediate applications. So big efforts has been made by the International Research community and industries in order to prove that, applying this technology, it’s possible to obtain joints with high mechanical properties and excellent fatigue behaviour, comparable to the base material. Actually Friction Stir Welding concretely demonstrated to be a really reliable joint technology, opening new possibility in designing simpler and efficient integral structures. After an initial tuning of the welding parameter, the process demonstrated to be really robust and repeatable. The presence of defects is clearly connected to imperfect process parameters, but the defect typology easy suggests appropriate corrections. What is more if a distinctive defect is present, it emerge in every point of the joint, so greatly simplifying inspection activity. In any case, on an optimized process, the new Non Destructive Inspection methods demonstrated to be really effective on monitoring crack propagation. Static and fatigue tests unquestionably verified joints high performance, also in aluminum alloy declared as difficult to weld. The presence of possible defect only slightly decrease the joint fatigue limit and high repeatability on the crack nucleation site was always observed. The reduced heat input allowed the manufacturing of structures with a limited degree of distortion and connected residual stresses. Simple prediction models based on a simple engineering approach that take into account only base material curves and experimentally measured residual stresses can correctly predict fatigue crack propagation, also in presence of complex load spectra. So Friction Stir Welding nowadays really represents the “Key technology” both for aeronautical and aerospace metallic structures.

41 SELECTED PAPERS I

4. SELECTED PAPERS

42 SELECTED PAPERS I

4.1 CONVENTIONAL STRUCTURES AND JOINING TECHNOLOGIES

4.1.1 BENDING

[1] A. Lanciotti, C. Polese “Fatigue crack propagation of through cracks in thin sheets under combined traction and bending stresses” Fatigue and Fracture of Engineering Materials and Structures , v. 26, 2003, pp. 421-428...... 44

[2] A. Lanciotti, C. Polese “Valutazione numerica della propagazione di difetti sotto l'azione combinata di carichi di trazione e flessione” XVII Congresso della Associazione Italiana di Aeronautica e Astronautica, Roma, 15-19 Settembre 2003...... 52

43 44

Fatigue crack propagation of through cracks in thin sheets under combined traction and bending stresses

A. LANCIOTTI and C. POLESE Department of Aerospace Engineering, University of Pisa, Via G. Caruso, 56122 Pisa, Italy Received in finalform 13 November 2002

ABSTRACT Fatigue crack propagation tests of through cracks in 3 mm thick aluminium alloy 6013- T6 plates, under combined membrane and bending stresses were carried out. Five different values for the ratio of bending to membrane stresses, SB, were examined: 0, 0.55, 1.25, 1.8 and 2.23. Firstly, the results were elaborated by taking into account only the membrane stress and the front dimension of the crack for the evaluation of the stress intensity factor range. The results relevant to the lower SB values, evaluated in this mode, show good agreement with the results obtained when only the membrane stress was applied, while the results obtained at the higher SB values exhibit a remarkable increase in the crack propagation rate. The results were successively evaluated on the basis of tabular stress intensity factor solutions available in the literature; the agreement between the predicted and the experimental data occurs when mean values of stress intensity factor, rather than local values, were used at the back face of the plate, i.e., the face where the bending produces a compressive stress, to counteract the high gradient of stress intensity factor present in this area. Keywords combined tension and bending stresses; experimental measurements; finite element analysis; stress intensity factor. NOMENCLATURE a ˆ crack depth c ˆ front surface dimension of the crack

ct ˆ back surface dimension of the crack e ˆ eccentricity of the specimen

fw ˆ finite width correction factor r ˆ radius of the hole t ˆ thickness y ˆ abscissa along the thickness of the plate K ˆ stress intensity factor

R ˆ stress ratio Smin/Smax SB ˆ secondary bending (bending stress/membrane stress) W ˆ width of the specimen b ˆ normalized stress intensity factor

joining technique in aircraft structures and surely is INTRODUCTION going to maintain this position for many years. Riveting Even if the fatigue behaviour of rivet joints has been is used to realize joints in different aircraft areas, but studied for more than 50 years, the nucleation and propa- recently the attention of researchers was particularly gation of fatigue cracks in these element are still the focused on fuselage riveted joints, also as a result of object of investigations. Riveting is the most popular recent catastrophic failures of lap-splice joints of trans- port aircrafts. Nowadays more sophisticated calculations can be performed to better understand the different Correspondence: A. Lanciotti. E-mail: [email protected] aspects of the fatigue resistance of riveted joints. Also,

ß 2003 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 26,421±428 421 422 A. LANCIOTTI and C. POLESE 45 riveting technology has been the object of experimental loading. As far as the knowledge of the present authors is investigations, which, among others, demonstrated that concerned, no experimental verification has been per- by increasing and controlling the force applied during formed on these analytical expressions. riveting, large improvements on fatigue life of riveted The aim of the present activity was the production of structures can be achieved.1 experimental results on fatigue crack propagation of Several factors contribute to the fatigue resistance of through cracks under combined membrane and bending riveted joints, and taking all these factors into account in stresses, and the evaluation of the results on the basis of the prediction models is a very difficult task, far away, for the stress intensity factor solutions available in the literature. moment, from its definitive solution; also for this reason, the common approach to this problem, such as to other DESIGN OF THE SPECIMEN AND TEST problems, is to investigate separately the different APPARATUS factors, hoping in the near future to be able to group them in a single model. One of these factors is the As in other investigations, a standard, single actuator, presence of a bending stress, superimposed on the mem- fatigue test machine applied the combined membrane brane stress. The bending stress is not necessary to and bending stresses. In Ref. [3], a specially designed transmit the load in a joint, as, on the contrary, the test fixture was described. Loads were transmitted to membrane stress is (membrane stress is used here to the plate by pins at the top and bottom grips. The desired indicate the stress, constant in the thickness of a plate, ratio between bending and membrane stresses was which is created when an axial load is applied to the obtained by varying the eccentricity between the middle plate). The bending stress is created as a consequence plane of the specimen and the loading line. A plate (for- of the inherent eccentricity of single lap or butt joints; cing plate) was connected to the grips to increase the the ratio between the bending and the membrane stresses bending stress in the test section. The two forcing plates is named `Secondary Bending', SB. Strain gauge meas- limited the test frequency to 2 Hz, in order to avoid urements of SB in longitudinal fuselage lap splices have dynamic effects. shown high values of this parameter, to the point that, A thin sheet specimen loaded under combined tension even if the main load condition was the pressurisation of and bending stresses was described in Fawaz and de the fuselage, even compressive strains were recorded near Rijck.6 The specimen is composed of three plates bonded lap splice joints.2 The bending stress increases the stress to form a butt joint with a large central gap, Fig. 1(a). on the fraying surfaces of riveted joints, so that small The central plate is the specimen while the external ones surface or corner cracks nucleate in this area, where are the grip plates. By properly designing the length of inspections are very difficult to perform. Subsequently, the three plates, a different ratio between bending and small cracks propagate in through thickness cracks that membrane strains can be obtained. The specimen can can be easily detected when they emerge out from the be designed by using simple formulas from the theory rivet heads. After this event, the residual life of the struc- of elasticity or by using finite element models. Anyway, ture is only a small part of its whole life, so that accurate the large out-of-plane displacement of the specimen, calculations are necessary for a reliable estimation of which reduces the ratio between bending and membrane safety margins and residual life of the cracked structure. stresses while the membrane stress is increased, must All the same, only few investigations have been carried be taken into account. A slightly different scheme was out on the behaviour of through cracks subjected to adopted in the present case. The specimen was con- combined traction and bending stresses, while surface nected to the grip plates by bolted double butt joints. cracks have been deeply investigated both from theoret- The eccentricity of the specimen, and consequently the ical and experimental points of view. bending to membrane stress ratio, can be changed by Some results on the propagation of through cracks in introducing four equal thickness spacers, as shown in thin sheets subjected to combined traction and bending Fig. 1(b). This scheme proved to be very effective, and stresses were presented in Phillips.3 Testswere carried out more than 30 tests were carried out by using the same at two different values of SB, 0.75 and 1.5. The results grip plates. obtained have shown an increase in the crack propagation As for the thickness of the specimens, it is difficult to rate when the bending stress was superimposed to the produce high SB values in thin sheets without reducing membrane stress in 2.3 mm thick sheets, while this effect too much the height of the specimen. For this reason, it was small in the case of 1 mm thick sheets. was decided to carry out the experimental activity on The stress intensity factor solution for a part-elliptical 3 mm thick specimens; the material was aluminium through crack nucleated from a hole is available.4, 5 alloy 6013-T6. The orientation of the specimens was LT. Different load conditions were considered: membrane A finite element model was used to establish the length stress, bending stress, biaxial membrane stress and rivet of the specimen and the grip plates; calculations

ß 2003 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 26,421±428 FATIGUE CRACK PROPAGATION 423 46

(a) (b) 3

0.5 Grip plate Grip plate 70 Specimen

Specimen Bonded joint 50

Spacer 3 Bolted joint 40

P 3

P

Fig. 1 (a) Edge view of the combined tension and bending 250 specimen proposed in Ref. [6], (b) specimen used in the present study. Fig. 2 Main dimensions of the specimen and elastic grip plates used in the present study.

60 was 250 mm; Fig. 2 shows the main dimensions of the specimen connected to the grip plates, together with the e =2mm 50 dimensions of the start notch introduced for the subse- quent crack propagation tests. A strain gauge instru- 40 mented specimen was used to experimentally verify the results obtained; Fig. 3 shows the comparison between 30 the predicted and the evaluated bending stresses in the centre of the specimen as a function of membrane stress, when an eccentricity of 2 mm was imposed on the speci- 20 Strain gauges FEM results men. It is evident from the figure that the behaviour of Bending stress, MPa the specimen can be accurately predicted. 10

0 TEST PROCEDURE 0 10203040506070More than 30 specimens were fatigue tested under con- Membrane stress, MPa stant amplitude loading. The ratio between the min- Fig. 3 Comparison between the calculated and the measured imum and the maximum membrane stress in the fatigue bending stresses in the test specimen. Eccentricity e ˆ 2 mm. cycles was R ˆ 0.1 in all tests. Note that as the bending stress is not proportional to the membrane stress, the performed suggested a free length of 40 mm for the stress ratio changes slightly when the bending stress is specimen and 70 mm for the grip plates. Grip plates also taken into account for its evaluation. were 3 mm thick too. These dimensions allowed us to Crack propagation tests were carried out under five obtain the higher SB value, 2.23, by inserting four 3 mm different SB values: 0, 0.55, 1.25, 1.8, and 2.23. The thick spacers, see Fig. 1(b). The width of the specimen crack propagation was monitored during the tests by an

ß 2003 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 26, 421±428 424 A. LANCIOTTI and C. POLESE 47 optical microscope (20x) mounted on a sliding guide 0.01 equipped with a digital display. Obviously, only the SB=0.55 external dimensions of the cracks could be measured SB=0 during the tests. Besides, the measurement of the back 1E-3 dimension of the crack demonstrated to be rather diffi- cult due to the presence of a necked, but unbroken, small ligament that rendered identification of the crack tip 1E-4 uncertain. To obtain additional information on the , mm/cycle N crack shape evolution, some tests were interrupted at /d c prefixed values of the crack length and the specimens d 1E-5 statically broken to allow the observation of the crack fronts. Crack fronts were measured in 31 equally spaced points along the thickness of the plates by means of an 1E-6 optical 100x stereo microscope equipped with an x-y 2 3 4 5 6 7 8 9 10 20 30 40 50 table with digital displays. The number of cycles that Membrane ∆K, MPa m had to be associated to each crack was established on the basis of crack propagation tests carried out on perfectly Fig. 4 Results of the crack propagation tests carried out at equal specimens, tested under the same conditions up to SB ˆ 0.55 compared with the results obtained at SB ˆ 0. final failure. It was assumed that the same external dimension was obtained at the same number of cycles. Acting in this mode, the scatter between the crack initi- 0.01 ation and crack propagation rates of different cracks was SB=1.25 suppressed. SB=0 1E-3

RESULTS AND DISCUSSION As a first attempt, to put the effect of bending stress into 1E-4 , mm/cycle evidence, the results obtained were evaluated in the form: N /d crack growth rate vs. stress intensity factor range, by c d taking into account only the membrane stress in the 1E-5 evaluation of the stress intensity factor; besides, the crack was considered as a through crack with straight front, having a dimension equal to the dimension meas- 1E-6 ured on the front side. The front side is defined as the 2 3 4 5 6 7 8 9 10 20 30 40 50 side where the bending produces a traction stress; the Membrane ∆K, MPa m back side is the opposite side. The presence of the hole in Fig. 5 Results of the crack propagation tests carried out at the centre of the specimens was taken into account, in SB ˆ 1.25 compared with the results obtained at SB ˆ 0. the evaluation of the stress intensity factor, by the analyt- ical expressions given in Kuo et al.7 This effect is not negligible when the crack is small compared to the hole behaviour is no longer present under the higher SB radius. values, 1.8 and 2.23; in these cases the crack propagation The results obtained are shown in Figs 4±7, compared rate evaluated as described previously is higher than in with the best fit curve of the results obtained in the case the case SB ˆ 0. SB ˆ 0. Different symbols indicate the results obtained in Figure 8 shows a typical crack front observed under different tests. The first two diagrams, relevant to the combined membrane and bending stresses; the bending lower SB values, 0.55 and 1.25, show results practically stress produces curved crack fronts, which can be ap- comparable with the results obtained in the tests carried proximated with an elliptical shape. This approximation out at SB ˆ 0; so, it can be concluded that the two is acceptable for long cracks, while small cracks can approximations introduced, straight crack front and deviate from this model, as observed also by Denman sole membrane stress acting, tend to counterbalance et al.8 However, by resigning to impose the crack shape, when SB is low. From the physical point of view, the it is not possible to dispose of stress intensity factor bending stress increases the front dimension of the crack solutions ready to be used. but the crack propagation rate is unaffected due to the The stress intensity factor of a part-elliptical through slowdown effect on the back side of the plate. This crack in a plate subjected to membrane and bending

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5 0.01 stresses or rivet loading was calculated by Fawaz, who used the Finite Element Method. With reference to the symbols shown in Fig. 9, the results obtained were- SB=1.8 SB=0 1E-3 given in tabular form in eight points along the thick- ness y/t ˆ 0.014, 0.058, 0.154, 0.360, 0.641, 0.846, 0.942 and 0.986. Calculations were performed for different a/c ratio in the range 0.2±2, a/t ratio in the range 1E-4

, mm/cycle 1.05±10 and r/t ratio in the range 0.5±2. N

/d To have available a more detailed description of b factor c d along the crack front of some of the cracks observed in 1E-5 the experimental activity, finite element calculations were performed by using the ABAQUScode. 9 The FEM model was a rectangular plate, width 254 mm, 1E-6 height 254 mm with a central hole. The stress intensity 2 3 4 5 6 7 8 910 20 30 40 50 factor was evaluated by the three-dimensional virtual Membrane ∆K, MPa m crack closure technique.10 This method uses the nodal Fig. 6 Results of the crack propagation tests carried out at SB ˆ 1.8 forces along the crack front and the nodal displacements compared with the results obtained at SB ˆ 0. behind the crack front to evaluate the local strain energy release rate and consequently the stress intensity factor. As an example, Fig. 10 shows the b factor in different points of the front of a crack having a/c ˆ 1, r/t ˆ 0.5, a/t ˆ 1.05 (see Fig. 9); the load condition is a membrane 0.01 stress. This crack is very similar to one of the cracks SB=2.23 actually observed in the tests carried out at SB ˆ 0.55. SB=0 1E-3

Crack front 1E-4

, mm/cycle y c t N /d c a d t 1E-5 x r c

1E-6 2 3 4 5 6 7 8 9 10 20 30 40 50 Membrane ∆ , MPa K m Fig. 9 Part-elliptical through crack. Fig. 7 Results of the crack propagation tests carried out at SB ˆ 2.23 compared with the results obtained at SB ˆ 0.

2.6 K b = I 2.4 s pa fw c t=7.80mm Present results Fawaz, Ref. [5] 2.2 SB=2.23 2.0 1mm 1.8 1.6 1.4 =3mm

t 1.2

1.0 1.0 0.8 0.6 Crack front 0.8 0.4 0.6 =13.76mm y/ t c 0.2 0.4 c/t 0.0 0.2

Fig. 8 Photograph of a crack observed in the tests carried out at Fig. 10 b factor along the front of a quarter elliptical through crack. SB ˆ 2.23. a/c ˆ 1, a/t ˆ 1.05 and r/t ˆ 0.5. Loading: membrane stress.

ß 2003 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 26, 421±428 426 A. LANCIOTTI and C. POLESE 49

The results obtained are practically coincident with the 2.4 results given in Ref. [5]. 2.2 The practical use of the numerical results requires pre- liminary discussion. Stress intensity factor increases rap- 2.0 idly approaching the back side of the plate. By assuming 1.8 that the crack maintains a quarter-elliptical shape, only SB two b-values must be used to predict the growth of the 1.6 dimension c and ct (see Fig. 9) of the crack. The choice of 1.4 the b-value used to predict the growth of the dimension 1.2 ct of the crack is not immediate and by using the value that b assumes on the back side of the plate will give rise 1.0 0.00 0.05 0.10 0.15 0.20 0.25 to an excessive growth of the dimension c of the crack. t c/W The use of mean stress intensity factor values, rather than local values, can solve this problem, because so doing the Fig. 11 Decrease of bending stress with crack propagation. crack growth at a point of the crack front is influenced also by the values that the stress intensity factor assumes SB=0.55 11 at the adjacent points. Looking once again at the results 80 given in Fig. 10, the mean b-value at the back side of the Experimental 70 plate can be calculated so that the new quarter-elliptical Predicted front produces an incremental cracked area equal to the 60 incremental cracked area evaluated by considering local 50 Sm=50MPa stress intensity factor values. In this particular case cal- , mm 40 c culation gives rise to a b factor equal to 1.238. This value 30 is in-between the values given in Ref. [5] at y/t ˆ 0.846 20 and y/t ˆ 0.942 (b ˆ 1.145 and b ˆ 1.377); not wanting to Sm=41.8MPa carry out again all the calculations described in Ref. [5], 10 the crack propagation was evaluated by using for the back 0 0 200000 400000 600000 800000 1000000 1200000 side of the plate the numerical values of b given in Cycles Ref. [5] at y/t ˆ 0.846. This is clearly a simplified mode of proceeding, but notwithstanding it gave rise to good Fig. 12 Comparison between the predicted and experimentally results as shown in the following. Interpolation between observed crack propagation curves (SB ˆ 0.55). the tabulated values was performed linearly.

Besides the finite width correction factor, the decrease SB=1.25 in bending stress due to the reduction in the stiffness of 80 the plate caused by crack propagation was also taken into 70 Experimental account, by FEM modelling a cracked specimen together Predicted with the grip plates; the ratio between bending and 60 membrane stresses at the crack tip was evaluated for 50 Sm=34.5MPa different crack lengths. The crack tip was substituted 40 , mm with a small notch having a radius r ˆ 0.1 mm, so as to c allow the evaluation of the stresses at the crack tip. 30 Comparative calculations performed with different 20 notch radii, 0.5 and 1 mm, gave rise practically to the Sm=31.4MPa 10 same results. As an example Fig. 11 shows the results obtained in the case SB ˆ 2.23. The bending stress de- 0 0 500000 1000000 1500000 2000000 creases practically linearly with the crack length and a Cycles crack having c/W ˆ 0.25 produces a 40% reduction of SB. This value decreased slightly for the lower SB values Fig. 13 Comparison between the predicted and experimentally observed crack propagation curves (SB ˆ 1.25). examined. The results obtained are shown in Figs 12±15 in the form of front surface crack dimension as a function of case, the crack propagated up to a value of a/c lower than number of cycles. The results are good for the lower 0.2, which is the minimum value analysed in Ref. [5]. values of SB, up to 1.8, while in the case SB ˆ 2.23 the Consequently calculations were performed by imposing fatigue life is overestimated by a factor 2. Actually, in this a/c ˆ 0.2. This imposition obviously has resulted in a

ß 2003 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 26,421±428 FATIGUE CRACK PROPAGATION 427 50

SB=1.8 negative effect; it must be however, considered that a 80 factor 2 in fatigue prediction is not a bad result.

70 Experimental As far as the crack shape is concerned, having evaluated Predicted the front and back dimensions of the cracks and the 60 hypothesis that the crack is a part elliptical through 50 crack, allow us to draw the crack front across the thick- ness of the plates and to compare these results with the 40 Sm=41.5MPa , mm

c crack fronts experimentally observed. Figures 16 and 17 30 show this comparison for the tests carried out under the 20 lowest and highest SB values, 0.55 and 2.23. In the first =30.7MPa Sm case the comparison is given at the same number of 10 cycles, while in the second case the comparison is per- 0 formed at the same `c ' dimension, because the number of 0 250 000 500 000 750 000 1 000 000 cycles was not accurately predicted as in the first case. Cycles The shape of the crack front was better predicted in the Fig. 14 Comparison between the predicted and experimentally second case, while in the first case the actual cracks were observed crack propagation curves (SB ˆ 1.8). straighter than those predicted.

CONCLUSIONS SB=2.23 More than 30 fatigue crack propagation tests on plate 50 specimens subjected to in phase membrane and bending Experimental stresses were performed. The bending stress was intro- 40 Predicted duced in the specimens by mounting them eccentrically with respect to elastic grip plates. The dimensions of the 30 specimens and grip plates were established on the basis , mm c 20 of a preliminary FEM calculation, whose results were in good agreement with the experimental ones obtained S =41.8MPa a/c<0.2 m 10 by a strain gauge instrumented specimen. Tests were carried out at five different values of bending to mem-

0 brane stress ratio, SB. Some tests were interrupted and 050000 100000 150000 200000 250000 the specimens broken in order to allow the measurement Cycles of the crack fronts. Fig. 15 Comparison between the predicted and experimentally To emphasize the effect of the bending stress, crack observed crack propagation curves (SB ˆ 2.23). growth rate results were compared with the results

Experimental SB=0.55 Predicted

N=0 N=120000 N=254500 N=390000 N=495000

3 2

, mm 1

Fig. 16 Comparison between the a 0 predicted and experimentally observed 0 5 10 15 20 25 30 35 40 45 crack fronts (SB ˆ 0.5). c, mm

Experimental SB=2.23 Predicted

N=0 N=34900 N=85000 N=115800 Np=237200

3 2

, mm 1 Fig. 17 Comparison between the a 0 predicted and experimentally observed 0 5 10 15 20 25 30 35 40 c, mm Np=79000 crack fronts (SB ˆ 2.23). Np=207000

ß 2003 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 26, 421±428 428 A. LANCIOTTI and C. POLESE 51 obtained under a pure membrane stress, by evaluating 3 Phillips, E. P. (1997) An experimental study of fatigue crack the stress intensity factor only on the base of the growth in aluminum sheet subjected to combined bending and membrane stress and the front dimension of the crack. membrane stresses. Nasa Technical Memorandum 4784. 4 Fawaz, S. A. (1998) Application of the virtual crack closure The results were fully comparable at the lower SB values, technique to calculate stress intensity factors for through cracks so it can be concluded that the two approximations with an elliptical crack front. Engng. Fract. Mech. 59, 327±342. introduced, straight crack front and sole membrane 5 Fawaz, S. A. (1999) Stress intensity factors for part-elliptical stress acting, tended to counterbalance. The crack propa- through cracks. Engng. Fract. Mech. 63, 209±226. gation rate evaluated as described previously was over- 6 Fawaz, S. A. and de Rijck, J. J. M. (1999) Thin-sheet, combined estimated at the higher SB values. tension and bending specimen. Exp Mech. 39, 171±176. The results were evaluated on the basis of the stress 7 Kuo, A., Yasgur, D. and Levy, M. (1986) Assessment of damage tolerance requirements and analysis, Vol. II, Analytical methods. intensity factor solution given in Ref. [5]. Good predic- Air Force Wright Aeronautical Laboratories, AFWAL-TR- tions of fatigue life and crack shapes were obtained by 86±3003. taking into account mean stress intensity factor values 8 Denman, D. E., Baldwin, J. D., Taylor, B. M. and Frey, M. J. rather than local values, due to the high gradient of the (2001) Short crack evolution in countersink geometries. In: stress intensity factor on the back face of the plate. Proceedings of the 21th Symposium of the InternationalCommittee on AeronauticalFatigue, 18±20 June, 2001. CeÂpadueÁs Edition, Toulouse. 9 ABAQUSVersion 6.2-001 , Copyrigth ß Hibbitt, Karlsson and REFERENCES Sorensen, Italy. 1MuÈller, R. P. G. and Hart-Smith, L. J. (1997) Making fuselage 10 Shivakumar, K. N., Tan, P. W. and Newman, J. C. Jr. (1988) riveted lap splices with 20-year crack-free fatigue lives. In: A virtual crack-closure technique for calculating stress intensity Proceedings of the 19th Symposium of the InternationalCommittee factors for cracked three dimensional bodies. Int. J. Fract. 36, on AeronauticalFatigue, 18±20 June, 1997. Engineering Materials R43±R50. Advisory Service, Edinburgh. 11 Cruse, T. and Besuner, P. (1975) Residual life prediction for 2 Yeong, D. Y., Roach, D. P., Canha, J. V., Brewer, J. C. and surface cracks in complex structural details. J. Aircraft 12, Flournoy, T. H. (1995) Strain fields in Boeing 737 fuselage lap 369±375. splices: field and laboratory measurements with analytical correlations. FAA Technical Center, DOT-VNTSC-FAA-95±10.

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Valutazione numerica della propagazione di difetti sotto l’azione combinata di carichi di trazione e flessione

Agostino Lanciotti, Claudia Polese Dipartimento di Ingegneria Aerospaziale, Università di Pisa Via G. Caruso, 56122 Pisa - Tel. 050 2217211, Fax 050 2217244 e-mail [email protected], [email protected]

Sommario

Nelle giunzioni rivettate di impiego aeronautico di tipo “lap splice”, l’eccentricità della linea di carico comporta l’introduzione nel giunto di una flessione, denominata “flessione secondaria”, che da luogo a sollecitazioni aggiuntive, non necessarie per trasmettere il carico. Presso il Dipartimento di Ingegneria Aerospaziale sono state condotte numerose prove di propagazione di difetti passanti in lastre di lega di alluminio 6013-T6 di spessore 3 mm, sottoposte a carichi combinati di trazione e flessione. I risultati sperimentali ottenuti hanno consentito di migliorare i modelli di previsione di crescita del difetto attualmente disponibili, che hanno mostrato limiti sia nella capacità di valutazione della velocità di propagazione, che nella previsione della forma del fronte del difetto. Mediante analisi FEM, eseguita con il codice ABAQUS, si è calcolato l’accrescimento di difetti passanti con fronte obliquo, valutando l’effetto della crescita della dimensione del difetto sulla rigidezza del provino e impiegando i valori locali del fattore K di intensificazione degli sforzi lungo il fronte della cricca. I risultati sono in buon accordo con i dati ottenuti sperimentalmente, sia in termini di numeri di cicli che di geometrie del fronte. L’analisi FEM risulta comunque un processo lungo e di difficile implementazione: si è quindi elaborato un modello semplificato di predizione di crescita che impiega valori tabulati del fattore K già presenti in letteratura. Mantenendo l’ipotesi di fronte semiellittico, si può valutare correttamente l’avanzamento del difetto calcolando la crescita solo per i due punti estremi del fronte. A questi punti devono essere associati opportuni valori medi del fattore K che tengano conto dell’evoluzione geometrica del fronte di cricca.

INTRODUZIONE

Sebbene il comportamento a fatica delle giunzioni rivettate sia oggetto di studio da più di 50 anni, la nucleazione e la propagazione di difetti per fatica in questi elementi sono tuttora investigate. La rivettatura infatti, è la tecnica di collegamento più utilizzata nelle varie aree del velivolo e sicuramente manterrà tale posizione ancora per molti anni; recentemente l’attenzione dei ricercatori si è prevalentemente concentrata sui giunti longitudinali di fusoliera, anche in seguito ad alcune rotture catastrofiche osservate. Oggigiorno possono essere condotti calcoli più sofisticati per comprendere meglio l’influenza dei vari fattori che determinano la resistenza a fatica di tali giunzioni. Raggruppare tali parametri in un modello unico di previsione risulta però un compito tutt’altro che facile, lontano per il momento dalla sua soluzione definitiva; di conseguenza l’approccio comune a questo problema è analizzarli separatamente. Uno dei fattori principali è la tensione flessionale, denominata “flessione secondaria”, che si genera nel giunto per effetto dell’eccentricità della linea di carico. Il rapporto tra la tensione flessionale e la tensione membranale introdotta è indicato come Kb: tale parametro risulta particolarmente utile per esprimere la severità degli sforzi flessionali quando questi sono accoppiati ad uno sforzo membranale noto. Rilevazioni estensimetriche del fattore Kb in giunti longitudinali di fusoliera [1] hanno mostrato alti valori di questo parametro a tal punto che, sebbene la condizione principale di carico sia la pressurizzazione della fusoliera, sono state registrate anche deformazioni in compressione vicino ai giunti a semplice sovrapposizione. La flessione secondaria da luogo a tensioni normali aggiuntive sul lato interno dei giunti favorendo la nucleazione di piccole cricche d’angolo o superficiali in quest’area dove è difficile condurre ispezioni; tali difetti, propagando, diventano passanti e possono facilmente essere individuati solamente quando emergono fuori dalla testa del rivetto. A questo punto la vita residua della struttura è solo una piccola porzione dell’intera vita; devono quindi essere condotti accurati calcoli per una stima efficace dei margini di sicurezza e della vita residua della struttura fessurata. Fino ad oggi sono stati condotti pochi studi sul comportamento a fatica di difetti passanti soggetti a carichi combinati di trazione e flessione, mentre i difetti superficiali sono stati analizzati approfonditamente, sia dal punto di vista teorico che sperimentale. Alcuni risultati sulla propagazione di difetti in queste condizioni di carico sono stati presentati da Phillips [2]. Le prove sono state condotte per due diversi valori

- 9 - 53 di Kb, 0.75 e 1.5. I risultati ottenuti mostrano un incremento nel rateo di propagazione di un difetto in lastre di 2.3 mm di spessore, mentre questo effetto è minore nel caso di lastre di 1 mm di spessore. Tramite analisi FEM condotte in [3] sono disponibili le soluzioni del fattore di intensificazione degli sforzi K per una cricca semiellittica che si nuclea da un foro. In tale studio sono state considerate diverse condizioni di carico: trazione, flessione, stato di tensione biassiale e carico dei rivetti. Allo stato attuale non erano state condotte verifiche sperimentali di tali espressioni analitiche. Lo scopo della presente attività è quindi l’esecuzione di una serie di prove sperimentali sulla propagazione per fatica di difetti sotto carichi combinati di trazione e flessione e la valutazione dei risultati sulla base delle soluzioni di fattori di intensificazione delle tensioni disponibili in letteratura e verificate nel presente lavoro.

PROGETTAZIONE DELL’ATTREZZATURA DI PROVA E DEI PROVINI IMPIEGATI

Per l’esecuzione delle prove è stato impiegato un attuatore servocontrollato della capacità massima di carico di 200 KN. I carichi di flessione e trazione sono stati introdotti nel provino realizzando un doppio collegamento bullonato con le piastre di afferraggio, secondo lo schema mostrato in fig. 1; il rapporto voluto tra tensioni di flessione e trazione è stato ottenuto variando l’eccentricità tra il piano mediano del provino e l’asse di carico tramite l’inserimento di opportuni spessori. Realizzando con il codice AFFERRAGGI GIUNTO BULLONATO ABAQUS [4] un modello agli elementi finiti, si sono fissate le P dimensioni necessarie per ottenere 5 differenti valori del PROVINO SPESSORI rapporto Kb, nel range 0 - 2.0: i calcoli suggeriscono una lunghezza di 40 mm per il Fig. 1 – Schem a dell’attrezzatura di prova. provino e di 70 mm per le piastre di afferraggio. Queste dimensioni consentono infatti di ottenere il valore più alto Kb, pari a 2.23, introducendo al massimo uno spessore di 3 mm.

Strumentando un provino con degli estensimetri 60 si sono verificati i risultati ottenuti: la fig. 2 e = 2 mm mostra il confronto tra i valori delle tensioni 50 )

flessionali calcolate e quelle rilevate al centro del a P

provino, in funzione della tensione membranale M 40 (

e applicata, quando viene imposta un’eccentricità l a n o di 2 mm. È evidente come il modello sia in grado i 30 s s

di predire adeguatamente il comportamento del e l f

e Letture estensimetriche provino. Questo schema di carico è risultato 20 n

o Risultati FEM i

particolarmente efficace e sono state così s n e

condotte più di 30 prove di propagazione. T 10

PROVE DI PROPAGAZIONE 0 0 10 20 30 40 50 60 70 Sono state condotte più di 30 prove di fatica Tensione membranale (MPa) per 5 diversi livelli di Kb, 0, 0.55, 1.25, 1.8 e 2.23, con carichi ad ampiezza costante e rapporto Fig. 2 - Confronto tra le tensioni flessionali calcolate e R tra tensione minima e massima pari a 0.1. La rilevate sperim entalm ente per e = 2 m m . propagazione del difetto è stata monitorata durante le prove mediante un microscopio ottico, 20x, montato su una slitta equipaggiata con un trasduttore di posizione dotato di display digitale.Ovviamente durante la prova si sono potute misurare solo le dimensioni esterne del difetto: in particolare è risultato difficoltoso anche valutare la dimensione esterna posteriore a causa della presenza di un sottile legamento, plasticizzato ma non rotto, presente soprattutto per alti valori del Kb. Per ottenere informazioni aggiuntive sull’evoluzione della forma del difetto, alcune prove sono state interrotte a prefissati valori di lunghezza del difetto ed i provini sono stati rotti staticamente per consentire un’accurata misura della geometria del fronte. I fronti sono stati rilevati in 31 punti nello spessore tramite un microscopio ottico, 100x, equipaggiato con 54 una tavola x-y con display digitali. Il numero di cicli da associarsi ad ogni lunghezza del difetto è stato stabilito sulla base di prove di propagazione condotte su provini identici, nelle stesse condizioni di carico, fino alla rottura finale, assumendo che si ottengano le stesse dimensioni esterne per lo stesso numero di cicli; in questo modo si elimina lo scatter esistente tra la nucleazione del difetto e le velocità di propagazione di cricche diverse.

ANALISI DEI RISULTATI

Per mettere in evidenza gli effetti della tensione flessionale, i risultati sono stati elaborati nella forma classica di velocità di propagazione del difetto in funzione della variazione nel ciclo del fattore di intensificazione delle tensioni, tenendo però in considerazione solo le tensioni membranali nel calcolo del
K. Inoltre la cricca è stata considerata come un difetto passante a fronte dritto, con dimensione pari alla lunghezza del difetto sulla faccia anteriore (la faccia dove la flessione produce una tensione di trazione, posteriore l’altra). I risultati ottenuti sono mostrati nelle fig. 3 - 6 e confrontati con la retta di best fit dei risultati ottenuti nel caso di Kb= 0.

0.01 0.01

Kb=0 Kb=0.55 Kb=1.25 Kb=0 1E-3 1E-3 ) ) o o l l c c i i c c / /

m 1E-4

m 1E-4 m m ( (

N N d d / / c c

d 1E-5

d 1E-5

1E-6 1E-6 2 3 4 5 6 7 8 9 10 20 30 40 50 2 3 4 5 6 7 8 910 20 30 40 50 0.5 0.5 ∆K membranale (MPa*m ) ∆K membranale (MPa*m )

Fig. 3 - Risultati di propagazione per Kb = 0.55. Fig. 4 - Risultati di propagazione per Kb = 1.25.

0.01 0.01

K =1.8 K =2.23 b Kb=0 b Kb=0

1E-3 1E-3 ) ) o o l l c c i i c c / /

m 1E-4

m 1E-4 m m ( (

N N d d / / c c

d 1E-5

d 1E-5

1E-6 1E-6 2 3 4 5 6 7 8 910 20 30 40 50 2 3 4 5 6 7 8 910 20 30 40 50 0.5 0.5 ∆K membranale (MPa*m ) ∆K membranale (MPa*m )

Fig. 5 - Risultati di propagazione per Kb = 1.8. Fig. 6 - Risultati di propagazione per Kb = 2.23.

I diversi simboli usati corrispondono a diverse prove. I primi due diagrammi, relativi ai valori più bassi di Kb, 0.55 e 1.25, mostrano risultati praticamente comparabili con il caso di tensione flessionale nulla; quindi si può concludere che le due approssimazioni introdotte, fronte del difetto dritto e sollecitazioni esclusivamente membranali, si controbilanciano quando la flessione secondaria è bassa. Da un punto di vista fisico la tensione flessionale aumenta la dimensione anteriore del difetto, ma il rateo di accrescimento non cambia a causa dell’effetto di rallentamento che si ha nel lato posteriore della piastra. Questo effetto non è 55 presente nel caso di alti valori di Kb, 1.8 e 2.23: in questo caso infatti, il rateo di accrescimento del difetto è più alto che nel caso di solo sforzo membranale. La fig. 7 mostra il tipico fronte di un difetto ct = 7.80 mm Kb=2.23 osservato sotto carichi combinati di trazione e flessione: la flessione produce fronti curvi, che 1 mm possono essere approssimati con un una forma

m ellittica. Questa approssimazione è accettabile m

3 per cricche lunghe, mentre difetti piccoli

=

t possono deviare da questo modello, come osservato in [5] e [6]. Rinunciando però ad assegnare la forma del difetto non è possibile disporre di soluzioni analitico-numeriche del c = 13.76 mm fattore di intensificazione degli sforzi K.

Fig. 7 - Foto di un difetto osservato sperim entalm ente per Il fattore K per difetti passanti in una piastra soggetta a carichi di trazione, flessione e carichi Kb = 2.23. dei rivetti sono stati calcolati tramite analisi agli elementi finiti da Fawaz [7]. Tali risultati sono stati implementati all’interno del codice di previsione di crescita del difetto Afgrow [8], sviluppato dall’Air Research Laboratory, nel modello di difetto passante con fronte obliquo, soggetto a carichi combinati. É stata quindi simulata con tale codice la propagazione di difetti rilevati sperimentalmente con cicli di carico sinusoidali con R = 0.1: come si può osservare nelle fig. 8 e 9 il codice sottostima ampiamente la vita a fatica rispetto a quanto riscontrato durante l’attività sperimentale.

Afgrow K costante 0.12 b Afgrow K costante 0.10 b Afgrow K variabile b Afgrow K variabile 0.10 b Dati sperimentali Dati sperimentali

] 0.08 ] m [ m

[ 0.08

a a c c c

i 0.06 c r i r C

0.06 C

a a z z z

0.04 z e i e i p 0.04 p m m A

0.02 A 0.02

0.00 0.00 0 100000 200000 300000 400000 500000 0 250000 500000 750000 1000000 1250000 Cicli Cicli Fig. 8 - Propagazione di un difetto per Kb = 0.55. Fig. 9 - Propagazione di un difetto per Kb = 1.25.

Si è quindi cercato di valutare gli effetti della diminuzione della tensione flessionale legata alla riduzione di rigidezza del provino utilizzato, durante la propagazione del difetto. A questo scopo è stato realizzato un modello agli elementi finiti, impiegando il codice ABAQUS, della struttura completa di afferraggio utilizzata durante le prove sperimentali di propagazione; per ogni configurazione di carico, si è calcolato il rapporto Kb all’apice del difetto per lunghezze del difetto c = 10, 20, 40 e 60 mm. Come si vede in fig. 10, il difetto è stato approssimato mediante una fessura passante a fronte Fig. 10 - Particolare della m esh all’apice del difetto nel m odello FEM . dritto posta nella mezzeria 56 del provino, perpendicolarmente all’asse di carico. L’apice del difetto è stato realizzato mediante un intaglio circolare di raggio r = 0.1; questa scelta permette di evitare singolarità nella soluzione numerica. Calcoli comparativi condotti con diversi valori del raggio di raccordo, 0.5 e 1 mm, hanno dato luogo praticamente

2.8 100 e=3 mm 2.4 e=2 mm e=1 mm e = 3 mm 2.0 80

1.6 e = 2 mm 60 % b

b K

1.2 K 40 0.8 e = 1 mm Kb 20 Kb % = ⋅100 0.4 Kb _ MAX

0.0 0 0 10 20 30 40 50 60 0 10 20 30 40 50 60 c [mm] c [mm] Fig. 11 - Variazione del Kb all’aum entare di c. Fig. 12 - Variazione percentuale del Kb all’aum entare di c. agli stessi risultati. Estraendo dall’output le tensioni all’apice dell’intaglio lungo la direzione del carico, si è ottenuto, come si osserva in fig. 11, un andamento decrescente del Kb all’aumentare della lunghezza della fessura; la propagazione del difetto provoca quindi una diminuzione della rigidezza flessionale e tale diminuzione risulta più accentuata per alte eccentricità. È interessante osservare però, fig. 12, come una fessura di 60 mm provochi una diminuzione di rigidezza di circa il 50% in tutti i casi analizzati. Questo risultato porta a concludere che percentualmente la variazione di rigidezza con l’avanzare del difetto non è influenzata dalla condizione di carico. Nota la diminuzione di rigidezza si è cercato di valutare come questa influisca sul calcolo della vita a fatica effettuato mediante un codice di previsione. Il codice Afgrow non prevede la possibilità di variare il rapporto tra tensioni flessionali e membranali durante l’avanzamento del difetto, pertanto le simulazioni precedenti sono state ripetute interrompendole al raggiungimento delle dimensioni del difetto c = 10, 20, 40 e 60 mm, e introducendo di volta in volta il valore effettivo di Kb relativo a quella dimensione del difetto. Le curve di propagazione sono state quindi confrontate con quelle fornite dal codice senza prevedere la variazione di rigidezza, e con quelle sperimentali. Si osserva però dai risultati riportati in fig. 9 per Kb = 1.25, come la variazione di rigidezza introduca solo un lieve miglioramento nel calcolo, dell’ordine del 10%. Si sono quindi analizzate più in dettaglio le soluzioni del fattore di Fronte cricca intensificazione degli sforzi K date in [7]: i

y ct risultati ottenuti sono riportati in forma tabulare per diversi rapporti geometrici del a

t fronte del difetto, fig. 13, a/c in un range da x 0.2 a 2, a/t tra 1.05 e 10 e r/t tra 0.5 e 2, e in r c otto punti lungo lo spessore, per y/t = 0.014, 0.058, 0.154, 0.360, 0.641, 0.846, 0.942 e 0.986. Per avere una descrizione più dettagliata del fattore K lungo il fronte del difetto sono state condotte analisi FEM Fig. 13 – Geom etria di un difetto passante con fronte ellittico. impiegando il codice ABAQUS. In analogia a quanto fatto in [9] il modello impiegato consiste di una piastra rettangolare di dimensioni 254 x 254 mm con un foro centrale e il fattore K è stato calcolato utilizzando la tecnica del Virtual Crack Closure (3D VCCT): tale metodo [10] utilizza le forze nodali lungo il fronte del difetto e gli spostamenti nodali davanti al fronte, per valutare localmente il tasso di rilascio energetico e di conseguenza il fattore K.. Questa tecnica è facilmente implementabile con il codice FEM utilizzato, inoltre, a differenza di altre, può essere impiegata anche per mesh con elementi disposti non ortogonalmente al fronte del difetto, come mostrato in fig. 14. Ciò risulta particolarmente vantaggioso quando, come in questo caso, il fronte del difetto interseca la superficie libera del provino. Come esempio si riporta in fig. 15 il confronto, in termini di fattore adimensionalizzato , 57 tra il calcolo del fattore K eseguito nel presente lavoro e i dati tabulati in [7] nel caso di un difetto sperimentalmente osservato per Kb = 0.55. Si è quindi effettuata la simulazione FEM, Fronte cricca tramite codice ABAQUS, della propagazione di un difetto in un provino soggetto ad una tensione membranale di 50 MPa con Kb pari a 0.55, di cui si hanno i dati sperimentali completi di propagazione. Inizialmente si è modellato sul bordo del foro un difetto passante la cui geometria è stata osservata sperimentalmente. Si è calcolato il fattore di intensificazione degli sforzi K per ogni elemento sul fronte del difetto utilizzando la tecnica del 3D VCCT, estraendo dall’output del programma i valori delle forze nodali sui nodi del fronte e gli spostamenti nodali sui nodi precedenti al fronte, Fig. 14 - M esh in prossim ità del fronte del difetto. nella direzione ortogonale al piano di frattura. Per gli elementi estremi del fronte il fattore K è stato calcolato nell’ipotesi di stato piano di tensione, per gli elementi interni nell’ipotesi di stato piano di deformazione, analogamente a quanto fatto in [3]. Ad ogni nodo sul fronte del difetto si è quindi associato un valore del fattore K come media 2.6 pesata dei valori nei due elementi contigui, sulla Risultati FEM Fawaz, Rif. [7] K base delle rispettive aree sul piano di frattura. β = 2.4 σ πaf Essendo la condizione di carico effettiva una W 2.2 combinazione nota di tensioni di trazione e 2.0 flessione, il fattore K è stato calcolato separatamente nei due casi, ottenendo il valore 1.8 complessivo per sovrapposizione. Per ogni nodo 1.6 il calcolo della propagazione è stato effettuato 1.4 utilizzando i risultati in termini di velocità di 1.2 propagazione ottenuti dalle prove condotte con 1.0 1.0 Kb = 0. Si è quindi valutata la propagazione del 0.8 fronte per passi successivi: il numero di cicli
N 0.6 0.8 0.4 0.6 per ogni passo è stato scelto in modo che la 0.2 0.4 crescita da associare ad ogni nodo non superasse 0.0 0.2 le dimensioni
di un elemento della mesh, nella Fig. 15 - Fattore 0 lungo il fronte di un difetto passante direzione di propagazione. Questa condizione permette inoltre di ritenere valida l’ipotesi che il con fronte ellittico; a/c = 1, a/t = 1.05 e r/t = 0.5. fattore K rimanga costante per ogni nodo in Carico solo m em branale. questo intervallo di propagazione. In fig. 16 si riportano due fronti calcolati per 35000 e 165000 cicli assieme ai due fronti di riferimento rilevati sperimentalmente per lo stesso numero di cicli e ai risultati di Afgrow. Si può notare come i calcoli FEM abbiano una buona concordanza con i dati sperimentali, sia in termini di vita a fatica che di previsione N=165000 cicli della geometria del fronte. Il Misure fronte calcolato da Afgrow si FEM Afgrow discosta molto da quello 3 sperimentale e risulta inoltre marcatamente meno obliquo.

] 2

m Ovviamente un tale processo m [

a 1 di previsione risulta troppo lungo e dispendioso in quanto è 0 1 2 3 4 5 6 7 8 9 necessario rimodellare il fronte c [mm] del difetto ad ogni step di calcolo. N=35000 cicli Si è quindi cercato di

Fig. 16 - Confronto tra i fronti calcolati, i dati Afgrow e il fronte rilevato implementare un modello di previsione semplificato che si sperim entalm ente per 6N pari a 35000 e 165000 cicli per Kb = 0.55. basa sugli stessi dati del fattore di 58 intensificazione degli sforzi K utilizzati dal codice Afgrow [7]. Assumendo che il fronte del difetto abbia una forma ellittica, è necessario calcolare l’avanzamento solo nei due punti estremi del fronte per ottenere le nuove dimensioni c e ct, fig. 13. Dall’analisi dell’andamento del K sul fronte del difetto si nota però, fig. 17, come esso presenti un picco sul lato posteriore del provino. 1.0 3 K Locale

0.8 K Locale K Medio K Medio 2

0.6 ]

m Fronte iniziale t m [ /

z

0.4 a 1

0.2

0.0 0 4 6 8 10 0 1 2 3 4 0.5 K [MPa(m) ] c [mm] Fig. 17 - Andam ento del fattore K lungo il fronte e relativa propagazione del difetto.

Utilizzando il valore locale del fattore K per la previsione di crescita della dimensione ct si ottiene una crescita eccessiva e un prematuro raddrizzamento del fronte del difetto. Questo è quello che sembra accadere all’interno del codice Afgrow, generando una stima di vita molto più bassa di quanto rilevato sperimentalmente. L’utilizzo di un valore medio, invece del valore locale, può risolvere questo problema: in questo modo, infatti, la crescita del punto estremo del fronte risulta influenzata anche dai valori del fattore K assunti nei punti adiacenti. A tal fine è stato scelto un valore tale che il nuovo fronte produca una variazione di area fessurata uguale a quella prodotta considerando il valore locale del fattore K. Tra i valori tabulati del fattore K, quello che più si avvicina al valore medio calcolato è quello per y/t = 0.856. É stata quindi realizzata una routine in codice Fortran che calcola l’avanzamento del fronte per passi successivi; ad ogni passo il nuovo valore del fattore di intensificazione degli sforzi K si determina interpolando linearmente tra i dati disponibili. La routine tiene conto anche della diminuzione delle tensioni flessionali causate dalla riduzione di rigidezza del provino. In questo modo sono state analizzate diverse condizioni di carico, partendo da geometrie del difetto rilevate sperimentalmente. Come si osserva dalle fig. 18 - 21 i risultati numerici sono molto buoni per Kb fino ad 1.8, mentre per Kb = 2.23 la vita fatica è sovrastimata di un fattore 2. In questo ultimo caso, però, il difetto, a causa dell’elevata tensione flessionale, assume configurazioni con a/c < 0.2, che rappresenta il valore minimo per cui si hanno disponibili i valori tabulati di K: il calcolo è stato quindi effettuato imponendo a/c = 0.2, con conseguente peggioramento della soluzione.

80 K =0.55 K =1.25 80 b Dati sperimentali b )

Risultati numerici ) 70 m

70 m m m

( Dati sperimentali (

60 c 60 c Risultati numerici 50 50 S = 34.5 MPa Sm = 50 MPa m 40 40

30 30

20 20 S = 31.4 MPa S = 41.8 MPa m 10 m 10

0 0 0 200000 400000 600000 800000 1000000 1200000 0 500000 1000000 1500000 2000000 Cicli Cicli Fig. 18 – Confronto tra la propagazione sperim entale Fig. 19 – Confronto tra la propagazione sperim entale

e stim ata di difetti per Kb = 0.55. e stim ata di difetti per Kb = 1.25.

59

80 ) 50 K = 1.8 K =2.23 b m b m )

70 (

m c 40

m Dati sperimentali

( 60 Dati sperimentali

c Risultati numerici Risultati numerici 50 30

40 Sm = 41.5 MPa 20 30 S = 41.8 MPa a/c<0.2 m 20 10 Sm = 30.7 MPa 10 0 0 0 50000 100000 150000 200000 250000 0 250000 500000 750000 1000000 Cicli Cicli Fig. 20 - Confronto tra la propagazione sperim entale Fig. 21 - Confronto tra la propagazione sperim entale

e stim ata di difetti per Kb = 1.8. e stim ata di difetti per Kb = 2.23.

Il confronto tra i fronti rilevati e quelli calcolati, riportati in fig. 22 e 23, per i valori di Kb = 0.55 e Kb = 2.23, mostrano un buon accordo dei risultati anche per quel che riguarda la forma del fronte del difetto.

Dati sperimentali K =0.55 b Risultati numerici

N=0 N=120000 N=254500 N=390000 N=495000

3 )

m 2 m (

1 a 0 0 5 10 15 20 25 30 35 40 45 c (mm)

Fig. 22 – Geom etrie del fronte del difetto stim ate e rilevate sperim entalm ente per Kb = 0.55.

Dati sperimentali K =2.23 b Risultati numerici

N=0 N=34900 N=85000 N=115800 Np=237200

) 3 m

m 2 (

a 1 0 0 5 10 15 20 25 30 35 40 c (mm) Np=79000 Np=207000

Fig. 23 - Geom etrie del fronte del difetto stim ate e rilevate sperim entalm ente per Kb = 2.23.

CONCLUSIONI

Sono state condotte più di 30 prove di propagazione di difetti su provini soggetti a carichi combinati di trazione e flessione. La flessione è stata introdotta nel provino montandolo in modo eccentrico rispetto alle piastre di afferraggio. Le dimensioni del provino e delle piastre di afferraggio sono state fissate sulla base di calcoli preliminari FEM, i cui risultati sono in accordo con quanto rilevato sperimentalmente tramite letture estensimetriche. Le prove sono state condotte per 5 diversi valori del rapporto Kb tra tensioni flessionali e membranali. Alcune prove sono state interrotte e i provini rotti staticamente per consentire la corretta misurazione della geometria del fronte del difetto. 60

Per enfatizzare gli effetti della tensione flessionale i risultati in termini di velocità di propagazione del difetto sono stati confrontati con i risultati ottenuti nel caso di sola tensione membranale calcolando il fattore di intensificazione degli sforzi K solo sulla base della tensione membranale e della dimensione esterna massima del difetto. I risultati sono praticamente equivalenti per i bassi valori di Kb per cui si può concludere che le due approssimazioni introdotte, fronte di cricca dritto e tensione solo membranale, tendono a bilanciarsi. Nel caso invece di alti valori di Kb il rateo di propagazione così calcolato risulta sovrastimato. I risultati sono stati valutati sulla base delle soluzioni del fattore di intensificazione degli sforzi K riportate in [7]. Si sono ottenute buone previsioni della vita a fatica e della forma del fronte del difetto facendo riferimento a valori medi del fattore K piuttosto che ai valori locali, a causa dell’alto gradiente del fattore K sul lato posteriore del provino e tenendo in considerazione anche la diminuzione delle tensioni flessionali durante l’avanzamento del difetto.

BIBLIOGRAFIA

[1] D. Y. Yeong, D. P. Roach, J. V. Canha, J. C. Brewer, T. H. Flournoy “Strain Fields in Boeing 737 Fuselage Lap Splices: Field and Laboratory Measurements with Analytical Correlations” FAA Technical Center, DOT- VNTSC-FAA-95-10, 1995. [2] E. P. Phillips “An Experimental Study of Fatigue Crack Growth in Aluminium Sheet Subjected to Combined Bending and Membrane Stresses” Nasa Technical Memorandum 4784, 1997. [3] S. A. Fawaz “Fatigue Crack Growth in Riveted Joints” Delft University Press, 1997. [4] ABAQUS Versione 6.3-001, Copyrigth © Hibbitt, Karlsson and Sorensen, Italia. [5] A. Lanciotti, C. Polese “Fatigue Crack Propagation of Through Cracks in Thin Sheets under Combined Traction and Bending Stresses” , Fatigue and Fracture of Engineering Materials and Structures, Vol. 26, N° 5, pp. 421- 428, 2003. [6] D. E. Denman, J. D. Baldwin, B. M. Taylor, M. J. Frey “Short Crack Evolution in Countersink Geometries” Proceedings of the 21th Symposium of the International Committee on Aeronautical Fatigue, 18-20 June 2001, Cépaduès Edition, Toulose. [7] S. A. Fawaz “Stress Intensity Factors for Part-elliptical Through Cracks” Engineering Fracture Mechanics Vol. 63, pp. 209-226, 1999. [8] Afgrow Version 4.0007.12.11, 2/19/03, Copyright © 1996-2003 AFRL/VASM. [9] S. A. Fawaz “Application of the Virtual Crack Closure Technique to Calculate Stress Intensity Factors for Through Cracks with an Elliptical Crack Front” Engineering Fracture Mechanics Vol. 59, N° 3, pp. 327-342, 1998. [10] K. N. Shivakumar, P. W. Tan, J. C. Newman Jr. “A Virtual Crack-Closure Technique for Calculating Stress Intensity Factors for Cracked Three Dimensional Bodies” International Journal of Fracture Vol. 36, R43-R50, 1988.

SELECTED PAPERS I

4.1.2 TECHNIQUES TO ENHANCE FATIGUE LIFE OF STRESSED HOLES

[4] A. Lanciotti, C. Polese “The effect of interference fit fasteners on the fatigue life of central hole specimens” Fatigue and Fracture of Engineering Materials and Structures, vol. 28, n. 8, 2005, pp. 587-597...... 62

[5] A. Lanciotti, C. Polese “Effetto dell'interferenza rivetto-foro sulla resistenza a fatica di provini con e senza trasferimento di carico" XVIII Congresso della Associazione Italiana di Aeronautica e Astronautica, Volterra, 19-22 Settembre 2005...... 73 [6] L. Boni, A. Lanciotti, R. Gonella, F. Pietra, C. Polese “Numerical simulations and experimental evaluations of some technological processes employed to improve the fatigue resistance of stressed holes” TCN-CAE 2005, International Conference on CAE and Computational Technologies for Industry, 5-8 October 2005, Lecce, Italy...... 84

[9] F. Franchini, F. Pietra, C. Polese, S. Urso Sculco “Chaboche Material Model Optimization for Evaluation of Residual Stress Field in Cold Worked Holes” International ModeFRONTIER Users’ Meeting, 28-29 September 2006, Trieste, Italy...... 90

[10] L.Boni, A. Lanciotti, C. Polese “Numerical and experimental assessment of fatigue behaviour of “cold worked” open holes”, Invited Lecture at the 17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy...... 109

[13] A. Lanciotti, F.Nigro, C.Polese “Caratterizzazione Damage Tolerance dell'attacco ala- fusoliera del velivolo M346” XVII Congresso della Associazione Italiana di Aeronautica e Astronautica, Roma, 15-19 Settembre 2003...... 130

[15] A. Lanciotti, F. Nigro, C. Polese “Damage Tolerance evaluation of the wing to fuselage connection of the new trainer aircraft M346”. Fatigue and Fracture of Engineering Materials and Structures, October 2006...... 141

61 62 doi: 10.1111/j.1460-2695.2005.00902.x

The effect of interference-fit fasteners on the fatigue life of central hole specimens

A. LANCIOTTI and C. POLESE Department of Aerospace Engineering, University of Pisa, Via G. Caruso, 56122 Pisa, Italy

Received in final form 25 November 2004

ABSTRACT Fatigue tests were carried out on 2024-T351, thickness 1.6 mm, central hole specimens containing pins installed with five different interference-fit levels. Tests clearly demon- strated the beneficial effect of interference fit on fatigue resistance, up to the maximum value examined, 2.5%. A three-dimensional (3D) finite-element model was used in or- der to characterize the stress field around the hole. A large specimen, with a 40-mm- diameter hole filled with interference-fit pin, was instrumented by strain gauges and statically tested in order to check FEM results. A very good correlation existed between measured and numerically evaluated strains. FEM results demonstrated the well-known effect of interference-fit fasteners on reducing stress ranges. By increasing the interfer- ence level, the stress range was practically unchanged, while the mean stress decreased. Interference-fit produces a biaxial stress state, which must be taken into account for fa- tigue evaluation. In the present case, a simple criterion, based on hoop strain, predicted the fatigue results quite well with the exception of open hole fatigue test results, which were overestimated. Keywords biaxial stresses; experimental measurements; finite-element analysis; interfer- ence fit.

NOMENCLATURE K f = fatigue stress concentration factor i = interference R = stress ratio = Smin/Smax S = nominal applied stress Seq = equivalent uniaxial stress Smin = minimum stress in fatigue cycle Smax = maximum stress in fatigue cycle Sx = radial stress Sy = longitudinal stress

fasteners, which are installed with the aid of a second ele- INTRODUCTION ment (a nut or a collar). When solid rivets are driven, the The fatigue resistance of mechanical joints in aircraft second head is formed but, at the same time, the shank structures is a problem not completely solved and continu- is also plastically deformed so that the hole is perfectly ously studied because, nearly always, the fatigue resistance filled, creating in this way a beneficial interference state. of aircraft structures corresponds to the fatigue resistance By increasing the force applied to drive rivets, interference of the joints contained in these structures. Limiting our increases too, and hence fatigue resistance. attention to mechanical joints, so many different elements Also threaded fasteners are often installed with interfer- are used that even their classification is difficult. A pos- ence, even if clearance is the usual installation standard. sible distinction can be made between solid rivets, which Ta pering is the common stratagem used to allow the in- are plastically deformed during installation, and threaded stallation of high interference fasteners; Smith1 quoted a U.S. patent dating back to 1886 for an adjustable ta- pered fastener by Henry A. Wahlert of the American Brake Correspondence: A. Lanciotti. E-mail: [email protected] Co., St. Louis, U.S.A.; although, the main purpose of this

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 587 588 A. LANCIOTTI and C. POLESE 63 system was to adjust for wear in brake parts and was not re- ference conditions, such as frictionless interface or fixed lated to fatigue. Also, cylindrical shank fasteners can be in- (no slip) interface, must be established. Under these hy- stalled with interference but high interference levels can- potheses, the problem can be also solved in an analytical not be obtained, as damages to the hole can be produced form.3,4 during installation. For this reason, special fasteners, such In the present paper, the simple case of an unloaded fas- as SLEEVbolt® (Gardena, CA), were developed; this fas- tener fitted in a hole (commonly named zero load transfer tener consists of a tapered bolt and an internally tapered joint) was examined; a lot of these joints exist in aircraft flanged outer sleeve, ground externally straight for use in structures, such as, for instance, the connection between straight-sided clearance holes. An interference fit, propor- longitudinal stiffeners and external fuselage skin, so this tional to bolt protrusion from the sleeve, is achieved as the simple case is of practical interest too. bolt head comes into contact with the structure. Fatigue tests were carried out on dogbone specimens The use of interference-fit fasteners can prove trou- having a central hole filled with different pins so as to ob- blesome, such as disassembling difficulties, which can tain five different interference levels, in the range from be solved, in joints that require frequent disassembling, 0.5 to 2.5%. Finite-element models were used to evaluate by installing clearance fasteners in interference-mounted the stress and strain fields around the hole when the fas- bushes. Further, excessive interference can promote stress teners were inserted and subsequently a fatigue load was corrosion cracking, particularly when the short transverse applied. The results obtained were the input for a simple grain direction is stressed. multiaxial fatigue criterion, which well correlates the re- Many tests have clearly demonstrated the beneficial ef- sults relevant to different interference-fit values, but was fect of interference fit on metallic structures’ fatigue re- not able to predict these results from the base material sistance; recent results have shown that beneficial effects fatigue curve. can be obtained also in composite materials.2 Despite this, although a lot of work has been carried out for other hole EXPERIMENTAL WORK fatigue-life-enhancement systems, such as cold working, there is little information on the quantitative evaluation Te sts were carried out on Alclad 2024-T351 aluminium of fatigue resistance of interference fitted holes. alloy specimens; this material, like all 2xxx series alloys, The problem is not simple; even if cyclic stresses in well- is extensively used in aircraft structures, particularly in designed fastener joints do not exceed, under normal load fuselage and lower wing skin construction. Specimen ge- conditions, the yield stress of the material, interference ometry and equipment are described in the following. can produce local plastic deformations. When an external load is applied, both the pin and hole deform; slippage Fatigue tests and loss of contact can take place. Moreover, a multi- axial state of stress is produced. Taking all these factors Fatigue tests were performed on 1.6-mm-thick dogbone into account in prediction models is not simple, so, as specimens, having a central reamed hole, diameter φ = explained previously, the beneficial effect of an interfer- 4.00 mm, Fig. 1; the holes were filled with different steel ence fit is well known, but a quantitative prediction of pins, to obtain five different interference values of 0.5, this effect is difficult to make. Also, the evaluation of the 0.75, 1.25, 1.875 and 2.5%. Steel, due to its higher elas- stress state around an interference-fit pin is not easy, unless ticity modulus, is more efficient than aluminium alloys by introducing simplifying hypotheses, such as elastic be- as a material to fill an open hole to improve its fatigue haviour in all conditions and no separation between sheet life. The main reason for the material selection was the and pin when an external load is applied. Besides, inter- possibility of performing all the fatigue tests having the

Fig. 1 Fatigue specimen.

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 64 THE EFFECT OF INTERFERENCE-FIT FASTENERS 589

Fig. 2 Specimen with hole filled by sleeve and pin.

Fig. 3 Strain gauge map in the large specimen used in static test. same interference level using the same pin, thus avoid- ing the additional source of scatter connected with the construction of nominally identical pins. Some specimens φ = 40 mm, Fig. 3, allowed us to measure strains by 16xy were tested without a pin, to dispose of comparative open strain gauges positioned near the hole along and across the hole fatigue test results. A sleeve was interposed in the load direction; the fatigue test specimens were too small contact between the hole and pin, to avoid damage to for this measurement. In this case too, a sleeve was inter- the hole when pins were inserted. FTI (Fatigue Tech- posed between the pin and the hole. After insertion of the nology Inc., Seattle, U.S.A.) sleeves, the same used for pin, the specimen was loaded by a 250 kN servo hydraulic cold working, were used for this purpose. Obviously, the test machine up to a maximum gross stress of 130.8 MPa sleeve thickness was taken into account in the evaluation and unloaded. Strains were recorded during the test. of the interference level between the pin and hole. The longitudinal cut in the sleeve was positioned in the load RESULTS AND DISCUSSION direction. Figure 2 shows a specimen with the hole filled by a sleeve and pin. Sleeves are self-lubricated so, even if The results of fatigue tests, expressed as the number of cy- the same pin was used to perform from 10 to 15 tests, no cles to final failure as a function of maximum applied stress, wear was measured. Additional tests were carried out using are shown in Fig. 4; tests were carried out for five differ- faceted pins; the details of these tests will be given later. ent interference-fit values. Additional tests were carried The ratio between minimum and maximum stresses in fa- out on open hole specimens (no pin) and plane specimens tigue cycles was R = Smin/Smax = 0.1 in all tests. Stresses (same geometry, without hole). The best-fit curve is given were evaluated with respect to the gross section of the for each group of specimens tested under the same con- specimens. ditions; a horizontal arrow indicated a run-out, i.e. a test To tally, more than 130 fatigue tests were carried out; a suspended, without fatigue damage evidence, at a number 10 kN servo hydraulic test machine was used to carry out of cycles greater than 4 × 106. The results obtained clearly fatigue tests and some tensile tests on small coupons to indicate the beneficial effect of interference on fatigue life obtain the stress–strain curve of the material. and this effect increases with the interference level; prob- ably the maximum beneficial effect was nearly obtained because the results relevant to the higher values tested, Static test on a large specimen 1.875 and 2.5%, are quite similar. Fatigue test results, described subsequently, were exam- In the same figure, the fatigue test results relevant to ined in terms of stresses and strains in the critical loca- specimens with a central cold worked hole are given too.5 tion, evaluated by finite-element method (FEM). A large The material, 2014-T6, was similar to the material used specimen, made of the same material, was built, to inves- in the present investigation; the stress ratio was R = 0.1 in tigate the actual capability of FEM models in predicting this case too and the thickness was comparable, t = 2 mm. the specimen strain field during the insertion of the pin Expansion during cold working was ∼4%. A comparison and subsequent loading and unloading. The large speci- between the two sets of results shows a better behaviour men, 200 mm wide, 6 mm thick and having a central hole of cold worked holes in comparison with interference-fit

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 590 A. LANCIOTTI and C. POLESE 65

350

300 Base material 250 i=1.25% i=0.75% i=1.875% 200 i=2.5% i=0.5% CW, [2] 150

ax (MPa) Open hole m S 100 Base material Open hole i=0.5% i=0.75% i=1.25% R=0.1 i=1.875% i=2.5% 50 105 106 107

Cycles Fig. 4 Fatigue test results. holes. Obviously, limited to the values of the parameters enough large to allow strain gauge installation around the used, focussing on a fixed number of cycles, 106,forin- hole. Interference was i = 1.25%. stance, the fatigue stress concentration factor (the ratio Strains were measured during the installation of the pin, between the allowable stress of unnotched and notched which was carried out with the aid of a hydraulic jack. specimens) is K f = 2.19 for the open hole, K f = 1.83 for Subsequently, the specimen was loaded and strains were the interference-fit filled hole (i = 2.5%) and K f = 1.63 measured at different applied loads. Some results are given for the cold worked hole. It is important to note that the in Fig. 5, where the longitudinal strain is plotted as a func- two systems, cold working and interference fit, can be si- tion of the distance from the hole edge, after the insertion multaneously applied to the same hole, probably giving in of the pin and at maximum stress, equal to 130.8 MPa. Ex- this case the maximum beneficial effect. In fact, the two perimentally measured strains are compared, in the same systems do not exclude each other, as they have a differ- figure, with the numerical results obtained by ABAQUS.9 ent effect on fatigue resistance; cold working introduces Calculations were carried out by using the experimentally beneficial compressive hoop stresses and work hardening, measured monotonic stress–strain curve of the material. while interference reduces stress ranges, even if it will be The FEM model had 8200 eight-node isoparametric solid shown that compressive hoop stresses are introduced also elements. Linear kinematic hardening behaviour was as- by high interference fit. sumed as this model quite well predict the behaviour of the Obviously, improving factors in fatigue resistance ob- material under examination.10 The contact between the tained by testing thin specimens containing a no load pin and hole was supposed to be frictionless. Interference transfer hole cannot be directly transferred to thicker was imposed by radially expanding the hole, while the pin specimens or to different specimen geometry, even if some was radially compressed. A good agreement between nu- results seem to indicate this tendency. For instance, ex- merical and experimental results was obtained, as shown periments carried out on 6.35 mm thick no transfer spec- in Fig. 5. The maximum strain at the hole edge is about imens6 and 4 mm thick low transfer specimens7 indicate 10 500 µε;this value is lower than the interference level, the beneficial effect of interference fit on fatigue life, while 1.25%, corresponding to 12 500 µε,asaconsequence of other results8 give the same indication for single-lap riv- the elasticity of the steel pin. eted specimens. Figure 6 shows longitudinal stresses acting in the spec- In order to interpret the fatigue results in terms of imen at different applied stresses, from pin insertion up stresses and strains in the critical location, we posed the to maximum stress. Plasticisation took place when the pin problem of a possible experimental explanation of results was inserted; the plasticized area increased when an exter- obtained by a FEM model of an interference-fit fastener. nal load was applied; maximum stress took place far from For this reason, a larger specimen was constructed; this the hole edge at a point which moved away when the load specimen is similar to the fatigue test specimens, but is was increased.

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 66 THE EFFECT OF INTERFERENCE-FIT FASTENERS 591

0.018 0.017 i=1.25% S 0.016 Experimental 0.015 FEM 0.014 0.013 S=130.8 MPa y 0.012 0.011 x 0.010 (mm/mm) 0.009 y ε 0.008 0.007 0.006 0.005 0.004 0.003 S=0 MPa 0.002 0.001 0.000 051015 20 x (mm)

Fig. 5 Strains recorded during loading of the large specimen.

350 i =1.25% S=130.8 MPa 300 S=114.4 MPa S= 81.7 MPa S= 49.0 MPa 250 S= 0.0 MPa

200

150 S (MPa) y

S

100 y x 50

0 Fig. 6 Longitudinal stress acting in the 01020304050 large specimen as a function of applied stress. x (mm)

The same calculations were made for the specimen used rial, taken from the literature,11 instead of the monotonic in fatigue tests; the model and the assumptions were sim- curve experimentally evaluated; the results obtained will ilar to those described previously for the large specimen. be discussed later. In each model, three different steps— Seven interference levels were examined, from 0 to 4%, pin insertion, application of the maximum stress and un- and three different applied stresses, 125, 150 and 170 MPa, loading to minimum stress (R = 0.1)—were examined. which covered the range of stresses were applied in fa- As an example, Fig. 7 shows the results obtained in the tigue tests. Intermediate results were obtained by inter- case: interference i = 1.25%, maximum applied stress 125 polation. Some comparative calculations were carried out MPa. Longitudinal stress is plotted as a function of dis- also by using the cyclic stress–strain curve of the mate- tance from the hole edge for the three load conditions.

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 592 A. LANCIOTTI and C. POLESE 67

350

300 Interference (i=1.25%) Interference + S=125 MPa Interference + S=12.5 MPa 250

S 200

(MPa) 150

y y

S x 100

50

0 02468 Fig. 7 Longitudinal stress during loading x (mm) and unloading of fatigue specimens.

300 S

200 y S 100 y x

0

-100 S=0 (MPa)

S -200 S=170 MPa S=17 MPa -300

-400 S x -500 Fig. 8 Longitudinal and radial stresses at 0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 the hole edge during loading and unloading of fatigue specimens (maximum applied Interference (%) stress S = 170 MPa).

Also in this case, plasticisation took place when the pin if this value was not present in the experimental activity. was inserted; additional plastic strains were introduced by The results show that even minimum interference level, applying the external load, so after unloading (R = 0.1), i = 0.5%, produces a considerable decrease in the longitu- the stress state was slightly changed. After this point, sub- dinal stress range in comparison with the open hole case. sequent fatigue cycles produce only elastic stresses. All the The decrease in stress range is mainly due to the increase results obtained under the same maximum stress (Smax = of minimum stress. Actually, this result was not experi- 170 MPa) are given in Fig. 8. Longitudinal and transver- mentally confirmed, because small beneficial effects in fa- sal (radial) stresses acting at the hole edge are plotted as tigue life were obtained with the lower interference level. afunction of interference level for three conditions: pin By increasing further interference level, the longitudinal insertion, loading at maximum stress, Smax = 170 MPa stress range remains practically constant, while the mean and unloading at minimum stress, Smin = 17 MPa. In- stress decreases continuously, up to an almost asymptotic terference i = 4% was included in the calculations, even value, which is reached at i = 4%. Due to plasticisation,

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 68 THE EFFECT OF INTERFERENCE-FIT FASTENERS 593

200

i=0.5% S 150

p

100 (MPa)

p

50

0 020406080100 120 Fig. 9 Pressure between pin and hole as a S (MPa) function of applied stress (i = 0.5%). longitudinal stress after pin insertion decreases with inter- sequent paper,13 the same author uses a biaxial fatigue ference level, reaching negative values when i > 2.5%. As parameter, based on alternating shear stress and maxi- for transverse stress, it has a small range in fatigue cycles mum normal stress acting on the shear stress plane, to while its absolute value progressively increases with inter- compare (not to predict) the fatigue behaviour of open ference level; obviously, transverse stress is not present in hole specimens and filled hole specimens. Stresses were the open hole case. evaluated by FEM models in several locations around Numerical calculations showed that the contact between the hole. The results obtained indicated that the most the pin and hole was maintained up to maximum stress severe stress state occurred at a location different from when the interference level was greater than 0.75%, while, the notch root. Fatigue tests, carried out at eight differ- in the case i = 0.5%, a loss of contact took place when ent degrees of interference, in the range from 0 to 7%, the applied stress was S = 116 MPa, as shown in Fig. 9, confirmed this observation, i.e. cracks nucleated at differ- where the pressure acting between the pin and hole in ent angles with respect to the horizontal position; fatigue the longitudinal direction is plotted as a function of the results show a strong increase of fatigue life with an inter- stress applied. After this point, the pin is not effective from ference level up to 1.7%. Above this value, a further in- the fatigue point of view and for this reason, in the case crease in the interference level had a small effect on fatigue i = 0.5%, the range of the longitudinal stress is slightly resistance. higher when compared with the values obtained under The same problem was approached in Ref. [6] by the higher interference levels, as shown in Fig. 8. finite-element method, for the stress state evaluation, in Interference-fit produces a biaxial stress state which must conjunction with the Coffin–Manson method to calculate be taken into account in a prediction model. Several ap- crack initiation life. Central hole specimens with 1.2% proaches were used for fatigue analysis of interference-fit interference-fit fasteners and cold worked hole specimens fasteners in holes. In Ref. [12] central hole specimens with were fatigue tested. Predictions were good for cold worked 1.6% interference-fit fasteners were fatigue tested; a sim- holes, while lives predicted were 50 times higher than ex- plified uniaxial approach, based on maximum alternating perimental ones in the case of interference-fit holes and normal stress, was used to interpret the results. Stresses, interference-fit pins installed in cold worked holes. evaluated by FEM models, reached maximum values at a In Ref.[7] the same problem was approached by the distance from the hole edge approximately equal to the Crossland multiaxial fatigue criterion,14 which is based on hole radius. These values were taken into account for fa- a linear combination of equivalent shear stress amplitude tigue analysis. Good predictions of fatigue crack initia- and maximum hydrostatic stress. The application of the tion life were obtained, but the experimental evidence is model to the fatigue test results obtained by low transfer of crack initiation at the hole edge, so the method seems joints with interference-fit fasteners (only one interfer- to be in contrast with the experimental evidence. In a sub- ence value, not specified) is very good.

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 594 A. LANCIOTTI and C. POLESE 69

Three different multiaxial fatigue methods were exam- the fatigue cycle are given by ined by Shatil and Page15:the maximum principal stress S = S − νS ; S = S − νS . range (coincident in this case with the maximum hoop eq min ymin xmin eq max ymax xmax 14 stress), the Crossland model and the McDiarmid shear Acting in this way plastic strains due to the pin insertion 16 stress parameter. The parameters were calculated both and first fatigue cycle are not considered in damage eval- at the notch root and around the hole, looking for their uation, even if plasticization is taken into account in the maximum values. Good correlations were obtained using evaluation of local stresses and strains. the maximum hoop stress and the Crossland parameter The direct comparison of the fatigue results cannot be evaluated at the notch root, while the best correlation for made when local stresses or strains are considered, be- the McDiarmid parameter was found by evaluating the pa- cause the actual stress ratio changes as a consequence of rameter at locations different than the notch root, which material plasticity and residual stresses after unloading. changed with the interference level. To allow this comparison, results were corrected using Other multiaxial fatigue criteria can be used to analyse the relationship given in Ref. [18] to evaluate the equiv- the fatigue behaviour of interference-fit fasteners but it alent maximum stress for the case R = 0ofafatigue test must be noticed that these criteria were mainly developed conducted on 2024-T3 material, with a stress ratio R = 0: to analyse the fatigue behaviour of steel under in-phase 0.56 or out-of-phase bending and torsion loads; with the same SR = Smax(1 − R) . aim, these methods are evaluated by checking their appli- cation to the few multiaxial fatigue test results available The same relationship was subsequently used to evalu- = for steel constructions.17 ate the equivalent maximum stress in the case R 0.1, From a practical point of view, the present problem seems in order to compare the results with those relevant to the simpler: the biaxial stress state is composed of a longitudi- base material. The results obtained are given in Fig. 10 = nal stress, whose variation constitutes the actual fatiguing where the maximum equivalent stress (R 0.1) is plotted stress and a transverse stress, whose value, at a fixed in- as a function of the applied stress and interference level. terference degree, is quite constant in fatigue cycles, as Equivalent stress changes in almost linear way with ap- shown previously in Fig. 8. Moreover, after the pin inser- plied stress and decreases with interference level. Also in tion and the first fatigue cycle, subsequent fatigue cycles, this diagram, it appears that an interference level greater at the same level, only cause elastic stresses. than 2.5% does not produce appreciable variation of the Wanting to remain in the field of simple methods, at least equivalent stress. Fatigue test results were interpreted on for a first analysis, hoop strain is the simpler and repre- the basis of the results given in Fig. 10, which allows the sentative parameter. The problem being elastic after the evaluation of the equivalent stress as a function of the first cycle, hoop strain can be converted into an equivalent applied stress and interference level. All fatigue test re- uniaxial stress, whose minimum and maximum values in sults obtained, interpreted in this way, are presented in

400

350

300 (MPa)

=0.1 250

R

Seq 200 Open hole 150 i =0.5% i =0.75% i =1.25% 100 i =1.875% i =2.5% 50 i =4%

0 120 130 140 150 160 170 180 Fig. 10 Equivalent stress as a function of S max (MPa) applied stress and interference level.

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 70 THE EFFECT OF INTERFERENCE-FIT FASTENERS 595

350 Open hole 300 Base material

(MPa) 250 i=0.5% i=0.75% =0.1

R i=1.25%

Seq 200

150 i=1.875% i=2.5%

100 105 106 107 Fig. 11 Fatigue test results evaluated on the basis of equivalent stress. Cycles

Fig. 11. The equivalent stress correlates very well the fa- Different approaches were also tested; Fig. 13 shows the tigue test results obtained under different interference results obtained by using the hoop stress range. The pa- levels, which are almost reduced to only one fatigue rameter used in Ref. [12] is very simple, but does not take curve. So, by inverting the process, knowledge of the fa- into account the biaxial stress state, so it is not able to cor- tigue curve relevant to an interference-fit value allows us relate the results. Even worse were the results obtained by to evaluate the results relevant to different interference the Crossland parameter, used in Refs [7,15]. The effect values. In contrast, fatigue life is underestimated with of compressive transverse stress on hydrostatic stress was respect to the base material fatigue curve, while fatigue so high that even negative values of the parameter were life of the open hole specimens is overestimated. As ex- obtained. plained previously, these results were obtained by us- Coming back to fatigue results, the effect of transverse ing the monotonic stress–strain curve of the material. compressive stress is negative from the fatigue point of Some comparative calculations were performed by us- view; some tests were carried out with faceted pins, shown ing the cyclic stress–strain curve. The numerical results in the inset in Fig. 14. Faceted pins or slotted holes obtained in this case were practically the same for the were examined as a system to prevent fretting in lug/fork filled hole specimens, while fatigue life of open hole spec- joints.19 In the present case, faceted pins were used to re- imens was overestimated in this case too, but by a lower alize a uniaxial stress state at the notch root. Fatigue tests amount. clearly demonstrated that the suppression of the trans- Coming back to the results of filled hole specimens, addi- verse stress increases fatigue life, as shown in Fig. 14, tional factors are present, such as fretting and the position where the results relevant to faceted and round pins, of the critical location, which was assumed to be at the i = 1.25%, are given. Besides, the fatigue critical location notch root. Previous results13 and the examination of the was again the notch root, as in the open hole case, while in fractured specimens show that the critical location moves the tests carried out with a round pin, under the same in- around the hole by increasing the interference level up to terference value, the critical location was rotated 39◦ with 40–50◦;even if the scatter in the results of crack initiation respect to the notch root, as shown previously in Fig. 12. around the holes is large, a clear tendency can be observed, The fatigue results relevant to the faceted pin, evaluated as shown in Fig. 12, where the results are plotted, with the in terms of equivalent uniaxial stress, would be placed in mean value of each group. So, for any criterion adopted, Fig. 11 slightly over the open hole fatigue results. calculations would have to be carried out in several lo- cations around the hole, looking for the critical value of CONCLUSIONS the parameter. Over and above this complication, the cor- rect modelling of the contact between the pin and hole, Fatigue tests were carried out on 2024-T351, thickness with no friction or no slip contact, becomes an important 1.6 mm, central hole specimens containing pins mounted parameter. with different interference-fit levels.

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 596 A. LANCIOTTI and C. POLESE 71

60

50 (degree)

θ 40

30

20 Fatigue crack θ 10

0 0.0 0.5 1.0 1.5 2.0 2.5 Fig. 12 Crack-initiation point as a function Interference (%) of interference level.

350 Open hole 300 i=0.5% Base material 250 i=0.75%

200 Max Hoop stress (MPa) i=1.25% 150 i=1.875%

i=2.5% 100 105 106 107 Fig. 13 Fatigue test results evaluated on the Cycles basis of maximum hoop stress.

Te sts clearly demonstrated the beneficial effect of inter- of the base material. Other multiaxial fatigue approaches, ference fit on fatigue resistance up to the maximum value such as maximum hoop stress and the Crossland parame- examined, equal to 2.5%. ter, gave rise to worse results. The finite-element method was used to evaluate the Crack nucleation took place at different angles from the stress and strain field around a hole after the insertion horizontal line as a function of the interference level, while of the pins and subsequent fatigue cycling. Hoop strain in open hole specimens the notch root was the critical at the notch root was used as a parameter to evaluate an location. equivalent stress; the fatigue results obtained under differ- Acknowledgements ent interference-fit levels practically reduced to only one fatigue curve when interpreted in this way, even if it was The authors gratefully acknowledge the contribution not possible to derive this curve from the fatigue curve given by the students Paola Apruzzese, Nicola Pasquale

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 72 THE EFFECT OF INTERFERENCE-FIT FASTENERS 597

250

i=1.25%; R=0.1 Round pin 200 Faceted pin

150 max (MPa)

S 4.05

φ

3.65 Load direction Faceted pin 100 Fig. 14 Fatigue test results of specimens 105 106 107 containing round or faceted pins (i = 1.25%). Cycles and Ing. Francesco Pietra in the performance of the ex- 9 ABAQUS Version 6.4, Copyrigth
c Hibbitt, Karlsson & perimental and numerical activities. Sorensen, Italy. 10 Poussard, C. G. C., Pavier, M. J. and Smith, D. J. (1994) Prediction of residual stresses in cold-worked fastener holes REFERENCES using the finite element method. American Society of Mechanical Engineers, Eng. Syst. Des. Anal. 64, 47–53. 1 Smith, C. R. (1965) Interference fit fastener for fatigue-life 11 Osgood, C. C. (1982) Fatigue Design. 2nd edn., Pergamon improvement. Exp. Mech. 5, 19A–23A. Press, Oxford. 2 Lanza DiScalea F., Cloud, G. L. and Cappello, F. (1998) Study 12 Crews, J. H., Jr. (1974) An elastoplastic analysis of a uniaxially on the effects of clearance and interference fits in a pin-loaded loaded sheet with an interference-fit bolt. NASA TN D-7748. cross-ply FGRP laminate. J. Composite Mater. 32, 783–802. 13 Crews, J. H., Jr. (1975) Analytical and experimental 3 Crews, J. H., Jr. (1972) An elastic analysis of stresses in a investigation of fatigue in a sheet specimen with an interference uniaxial loaded sheet containing an interference-fit bolt. NASA bolt. NASA TN D-7926. TN D-6955. 14 Crossland, B. (1956) Effect of a large hydrostatic pressure on 4 Hou, J. P. and Hills, D. A. (2001) Interference contact between the torsional fatigue strength of an alloy steel. Proc. Int. a pin and plate with a hole. J. Strain Anal. Engng. Des. 36, Conference on the Fatigue of Metals. The Institution of 499–506. Mechanical Engineers, London, pp.138–149. 5 Cook, R. (1995) Fatigue enhancements due to cold expansion in 15 Shatil, G. and Page, A. G. (2002) Multiaxial fatigue analysis of a range of aluminium alloy. International Committee on interference-fit steel fasteners in aluminum Al 2024-T3 Aeronautical Fatigue, 18th Symposium, Melbourne, Australia, specimens. American Society Testing Materials, Special Technical 3–5 May, pp. 971–986. Publication, pp. 192–208. 6 Larouche, S., Bernard, M., Bui-Quoc, T., Julien, D., et al. 16 McDiarmid D. L. (1994) Shear stress based critical-plane (2001) Influence of cold working and interference fit on fatigue criterion of multiaxial fatigue failure for design and life life of 7475-T7351 aluminium alloy fastener hole. International prediction. Fatigue Fract.Engng Mater. Struct. 17, Committee on Aeronautical Fatigue, 21st Symposium, 1475–1484. Toulouse, 27–29 June, pp. 681–698. 17 Wang, Y. Y. and Yao, W. X. (2004) Evaluation and comparison 7 Duprat, D., Campassens, D., Balzano, M., Boudet, R., et al. of several multiaxial fatigue criteria. Int. J. Fatigue 26, (1996) Fatigue life prediction of interference fit fastener and 17–25. cold worked holes. Int. J. Fatigue 18, 515–521. 18 Anon. (1998) Metallic materials and elements for aerospace 8Muller,¨ R. P. G. (1995) An experimental and analytical vehicle structures. Military Handbook, MIL-HDBK-5H investigation on the fatigue behaviour of fuselage riveted lap 19 Schijve, J. (2001) Fatigue of Structures and Materials. Kluwer joint. Ph.D. Thesis, Delft University of Technology. Academic Publishers, Dordrecht, The Netherlands.

c 2005 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 28, 587–597 73

XVIII CONGRESSO NAZIONALE AIDAA 19-22 settembre 2005 VOLTERRA (PI)

EFFETTO DELL'INTERFERENZA PERNO/FORO SULLA RESISTENZA A FATICA DI PROVINI CON E SENZA TRASFERIMENTO DI CARICO

A. LANCIOTTI, C. POLESE

Dipartimento di Ingegneria Aerospaziale, Università di Pisa, Pisa

SOMMARIO

Si descrivono i risultati di un programma di prove di fatica effettuate su due diversi tipi di provini in lega leggera 2024-T351; entrambi i provini presentano un foro nel quale è installato un perno con interferenza e si differenziano per il carico trasferito dal perno, pari allo 0% e al 100% del carico applicato. Le prove di fatica hanno mostrato l'effetto benefico dell'interferenza sulla resistenza a fatica, fino al valore massimo esaminato, pari a 2.5%. Con analisi agli elementi finiti si è valutato lo stato di tensione intorno al foro. L'interferenza produce uno stato di tensione biassiale, composto da una tensione circonferenziale fortemente variabile nel ciclo di fatica e da una tensione radiale quasi costante. Per tener conto della tensione radiale sulla resistenza a fatica si è introdotta una tensione circonferenziale equivalente che fornisce, in senso circonferenziale, la stessa deformazione prodotta dallo stato di tensione biassiale. I risultati delle prove di fatica, interpretati con la tensione equivalente, vengono praticamente ridotti ad una sola curva, ma tale curva non è la stessa per i due diversi provini. Il metodo non appare quindi generalizzabile.

1. INTRODUZIONE

La resistenza a fatica delle giunzioni presenti nelle strutture dei velivoli è ancora oggetto di numerose indagini. Limitando l'attenzione alle giunzioni meccaniche, talmente numerosi sono gli organi di collegamento utilizzati, che anche la loro classificazione risulta complessa. Una possibile distinzione può essere effettuata fra "rivetto solido", costituito da un solo elemento che viene deformato plasticamente durante l'installazione e "rivetto composto", la cui installazione avviene con l'ausilio di un secondo elemento. Tale classificazione, anche se lacunosa, è sufficiente per la discussione che segue. Durante la ribattitura di un rivetto solido non solo si ha la formazione della seconda testa, ma anche il gambo viene deformato plasticamente, riempiendo perfettamente il foro. Si genera in questo modo un'interferenza fra foro e rivetto, che risulta essere benefica dal punto di vista della resistenza a fatica. Deformando maggiormente il rivetto si produce, oltre l'ovvio aumento della dimensione della testa, anche un aumento dell'interferenza e quindi della resistenza a fatica. Anche i rivetti filettati possono essere montati con interferenza, ma il montaggio con gioco è il più comune poiché il danneggiamento prodotto nel foro durante l'installazione può superare l'effetto benefico dell'interferenza. A tale proposito esistono rivetti speciali, quali lo SLEEVbolt® (Gardena, CA), che è costituito da un perno conico e da una bussola internamente conica ed esternamente cilindrica. In questo modo è possibile ottenere interferenza elevata senza danno per il foro. L'installazione di un rivetto con interferenza presenta anche alcuni aspetti negativi, quali difficoltà nello smontaggio e possibilità di rotture per tenso-corrosione, particolarmente quando viene sollecitata la direzione trasversale corta di un materiale. Anche se moltissime prove sperimentali hanno chiaramente dimostrato l'effetto benefico dell'interferenza sulla resistenza a fatica, non è ancora possibile prevedere tale effetto in modo quantitativo a causa delle complicazioni 74 introdotte da diversi fattori quali: possibili deformazioni in campo elasto-plastico, ulteriore plasticizzazione conseguente all'applicazione dei carichi esterni, triassialità dello stato di tensione, possibile scorrimento del rivetto rispetto al foro, unitamente a possibile distacco. Solo imponendo alcune ipotesi semplificative, quali comportamento elastico, assenza di separazione, assenza di attrito oppure assenza di scorrimento fra rivetto e foro, il problema è risolubile anche in forma chiusa, [1, 2]. Negli altri casi, che sono poi quelli di pratico interesse, è necessario ricorrere a modelli numerici. In una attività precedente è stato analizzato il caso di un rivetto che non trasferisce carico, [3]. A questo schema sono riconducibili molte giunzioni presenti nelle strutture dei velivoli, quali ad esempio le giunzioni fra irrigidimenti longitudinali e rivestimento di fusoliera. Nel presente lavoro, dopo aver brevemente riassunto i risultati dell'attività precedente, si esamina il caso di rivetto installato con interferenza che trasmette carico per taglio. Volutamente è assente un pretensionamento assiale del rivetto, che consentirebbe la trasmissione del carico anche per attrito. Questo schema è di interesse pratico per i collegamenti nei quali deve essere consentita la rotazione del giunto. Sviluppi futuri dell'attività riguarderanno il caso di giunzione che trasmette carico per taglio e per attrito.

2. PROVINO SENZA TRASFERIMENTO DI CARICO

2.1 Descrizione del provino

Il provino senza trasferimento di carico è un provino tipo "dogbone" con foro centrale. Il materiale utilizzato è lega di alluminio 2024-T351. spessore 1.6 mm, fig. 1. Il provino presenta un foro centrale alesato, φ=4.00 mm, nel quale sono stati installati diversi perni in modo da ottenere cinque diversi valori dell'interferenza, 0.5, 0.75, 1.25, 1.875 e 2.5%. I perni sono stati realizzati in acciaio; ciò, oltre che essere preferibile dal punto di vista della resistenza a fatica per il maggior modulo elastico del materiale, ha anche consentito l'uso di un solo perno per l'esecuzione di tutte le prove effettuate con lo stesso livello di interferenza. In questo modo si è eliminato l'ulteriore fattore di dispersione dei risultati connesso alla costruzione di diversi perni nominalmente identici. Per completezza sono state effettuate anche alcune prove comparative su provini di metallo base e su provini con solo foro centrale. Quest'ultima geometria viene denominata "foro aperto" per distinguerla dal caso di foro con installato un rivetto, denominata "foro chiuso". Per non danneggiare i fori durante l'installazione dei perni, si sono utilizzate le bussole elastiche prodotte dalla FTI (Fatigue Technology Inc., Seattle) per il "Cold Working" dei fori. Ovviamente si è tenuto conto dello spessore della bussola, pari a 0.152 mm, nel valutare l'interferenza fra foro e perno. Per mantenere la simmetria nel provino, il taglio longitudinale della bussola è stato orientato nella direzione del carico. La fig. 2 mostra il dettaglio di un perno installato in un foro con interposta la bussola elastica. Le bussole sono auto-lubrificate, per cui anche utilizzando lo stesso perno in 10÷15 prove di fatica, non si è rilevata usura. Ulteriori prove sono state effettuate utilizzando perni parzialmente spianati, come illustrato successivamente. In tutte le prove il rapporto di sollecitazione era R=Smin/Smax=0.1. Le tensioni sono riferite all'area lorda della sezione ristretta del provino (1.6x20 mm).

2.2 Risultati delle prove e discussione

I risultati delle prove di fatica sono riportati nella fig. 3, unitamente alla curva di best fit di ciascun gruppo (una freccia orizzontale indica un run-out, ossia una prova sospesa dopo un elevato numero di cicli senza evidenza di danneggiamento). I risultati chiaramente indicano l'effetto benefico dell'interferenza sulla resistenza a fatica, fino al valore massimo esaminato, pari a 2.5%, per il quale si osserva una "saturazione" della resistenza a fatica con l'interferenza. Per poter interpretare i risultati delle prove di fatica si è effettuata la modellizzazione agli elementi finiti del provino con il programma ABAQUS, [4]; il modello è composto da circa 8200 elementi isoparametrici a otto nodi. Per il materiale si è assunto il modello di incrudimento cinematico, che in altre sperimentazioni si è mostrato il più adatto per il materiale in oggetto, [5]. La curva tensione/deformazione (monotona) del materiale è stata determinata sperimentalmente e per i calcoli si è utilizzata questa curva invece della curva ciclica; tale approssimazione è accettabile poiché dopo l'introduzione di un perno con interferenza, le tensioni cicliche si mantengono in campo elastico. Si è ipotizzata assenza di attrito nel contatto perno/foro; l'interferenza è stata simulata espandendo radialmente il foro e comprimendo radialmente il perno. I calcoli sono stati ripetuti per sette diversi valori dell'interferenza, nell'intervallo 0-4% e per tre diversi valori della tensione applicata, 125, 150 and 170 MPa, che coprono l'intervallo delle sollecitazioni utilizzate nelle prove di fatica. I risultati intermedi sono stati interpolati. Le fig. 4 e 5 mostrano lo stato di deformazione e di sollecitazione in un provino conseguenti all'inserimento di un perno con bassa interferenza, 0.5 %, e con alta 75 interferenza, 1.25%. Nel primo caso lo stato di sollecitazione è praticamente elastico; tutte le grandezze (tensioni, deformazioni e tensione equivalente secondo Von Mises) raggiungono il valore massimo sul bordo del foro (il sistema di riferimento è definito nell'inserto contenuto nella fig. 6). Nel caso di interferenza elevata la tensione circonferenziale, Sy, raggiunge il valore massimo ad una certa distanza dal bordo del foro, essendo presente plasticizzazione. Il valore raggiunto sul bordo del foro è inferiore alla tensione di snervamento del materiale e tale valore diminuisce aumentando ulteriormente il livello di interferenza, fino a raggiungere valori negativi. Ciò può essere spiegato anche qualitativamente: mentre per la tensione circonferenziale è possibile una ridistribuzione, la tensione radiale è diretta conseguenza della pressione esercitata dal perno. Lo stato di tensione è fortemente biassiale e la tensione equivalente rappresenta un limite che non può essere superato. Di conseguenza aumentando la tensione radiale si ha una diminuzione della tensione circonferenziale. In campo elasto-plastico solo le deformazioni seguono l'andamento che intuitivamente si attribuirebbe alla soluzione elastica del problema, fig. 4 e 5. Applicando un carico esterno si manifesta una ulteriore plasticizzazione; come esempio la fig. 6 mostra il caso di un provino avente un perno con interferenza, i=1.25%, al quale viene applicato una tensione affaticante, Smax=150 MPa, R=0.1. La zona plasticizzata si allarga, come rilevabile dall'allontanamento dal bordo del foro del punto sottoposto alla sollecitazione massima. Scendendo al carico minino il diagramma di sollecitazione trasla semplicemente verso il basso; da questo punto in poi i cicli di fatica producono sollecitazioni elastiche. La fig. 7 mostra la sintesi dei risultati ottenuti per il livello di sollecitazione S=170 MPa. In funzione dell'interferenza vengono forniti i valori della tensione circonferenziale e della tensione radiale sul bordo del foro conseguenti all'introduzione del perno, alla successiva applicazione di una sollecitazione di 170 MPa e allo scarico fino a 17 MPa, (R=0.1). E' evidente la riduzione dell'escursione di sollecitazione circonferenziale conseguente all'introduzione di un perno con interferenza. Oltre il valore i=0.75% l'escursione di sollecitazione si mantiene costante, mentre diminuisce il valor medio della sollecitazione. Per i=0.5% l'escursione di sollecitazione è leggermente più ampia poiché dai calcoli risulta il distacco del perno dal foro, quando la sollecitazione raggiunge il valore S=116 MPa. Oltre il valore i=0.75% il perno mantiene il contatto con il foro fino alla sollecitazione massima. Per quanto riguarda la tensione radiale, questa aumenta con l'interferenza, ma la sua escursione nel ciclo di fatica è relativamente piccola. Diverse metodologie sono state utilizzate per la previsione della vita a fatica di elementi contenenti organi con interferenza. In [6] viene utilizzato il metodo degli elementi finiti, unitamente alla legge di Coffin-Menson per la previsione della vita a fatica di provini contenenti un perno con interferenza i=1.2%. I risultati sono buoni nel caso di foro trattato con Cold Working, mentre nel caso di foro non trattato la vita media stimata è circa 50 volte più alta della vita osservata in prova. In [7] lo stesso problema viene affrontato con il criterio di fatica multiassiale di Crossland [8]. I risultati ottenuti per un giunto a basso trasferimento di carico, con rivetti montati con interferenza (un solo livello di interferenza, non specificato) sono buoni. In [9] il problema viene affrontato con tre diversi metodi: l'escursione della tensione principale massima, il modello di Crossland [8] e il modello di McDiarmid, [10]. I parametri dei diversi modelli vengono calcolati in diverse locazioni intorno al foro, ricercando il loro valore massimo. I risultati sono buoni utilizzando la tensione principale massima e il modello di Crossland alla radice dell'intaglio, mentre per il metodo di McDiarmid i migliori risultati si ottengono valutando i parametri in posizioni diverse dalla radice dell'intaglio, che cambiano con il livello di interferenza. L'osservazione sperimentale in effetti mostra che in presenza di interferenza il punto critico non è alla radice dell'intaglio, [12, 3]. Dal punto di vista pratico, lo stato di tensione biassiale nel caso in esame è composto dalla tensione circonferenziale, la cui variazione costituisce l'elemento realmente affaticante e dalla tensione radiale che, fissato il valore dell'interferenza, si mantiene quasi costante in un ciclo di fatica, come già mostrato in fig. 7. Inoltre dopo l'inserzione del perno ed il primo di ciclo di fatica, i cicli successivi producono solo tensioni in campo elastico. Volendo rimanere nell'ambito dei metodi semplici, si osserva che l'effetto della pressione radiale è quello di produrre una deformazione circonferenziale legata, tramite il modulo di Poisson, alla deformazione radiale. Si può immaginare la stessa deformazione circonferenziale come prodotta da una tensione circonferenziale fittizia, il cui valore è uguale al prodotto della tensione radiale per il modulo di Poisson. Questa tensione sommata alla tensione circonferenziale fornisce una tensione equivalente monoassiale, con la quale si è valutata la resistenza a fatica. I valori minimo e massimo nel ciclo di fatica della tensione equivalente sono quindi dati dalle relazioni:

S =S -νS ; S =S -νS eq min ymin xmin eq max ymax xmax

76

Agendo in questo modo la deformazione plastica conseguente alla inserzione del perno ed al primo ciclo di fatica non viene considerata, anche se la plasticizzazione è tenuta in conto nella valutazione delle tensioni e delle deformazioni locali. Esaminando i risultati in termini di tensioni e deformazioni locali, sorge il problema del rapporto di sollecitazione che viene alterato rispetto al valore nominale. Per ricondurre tutti i risultati allo stesso valore del rapporto di sollecitazione, R=0.1, si è utilizzata la relazione fornita in [11], che consente di valutare la tensione massima equivalente per il caso R=0 per prove di fatica condotte sulla lega 2024-T3, con rapporti di sollecitazione R≠0:

S=S (1-R)0.56 R max

La stessa relazione è stata poi utilizzata per valutare la tensione equivalente per il caso R=0.1, in modo da confrontare i risultati ottenuti con quelli relativi al materiale base. I risultati ottenuti sono riportati in fig. 8; la tensione equivalente correla bene i risultati delle prove di fatica effettuate con valori dell'interferenza diversi da zero. Invertendo il processo, è quindi possibile prevedere, noti i risultati relativi ad un livello di interferenza, i risultati relativi ad altri livelli di interferenza. La vita a fatica non è comunque prevedibile partendo dalla curva di fatica del materiale base. Il caso di foro aperto è evidentemente diverso; in questo caso lo stato di tensione è monoassiale per cui i risultati, essendo analizzati in termini di tensioni e deformazioni locali, sono ragionevolmente deducibili dalla curva di fatica del materiale base. L'esame delle superfici di frattura ha mostrato, come già in altre sperimentazioni, [12] che mentre per il caso di foro aperto la sezione critica è la mezzeria del foro, nel caso di perno con interferenza il punto critico si sposta lungo il foro fino ad un angolo di 40-45°. Volendo migliorare le previsioni sarebbe quindi necessario valutare i parametri in diversi punti intorno al foro, ricercando la situazione più critica. Ciò implica un notevole appesantimento del lavoro di previsione. A questo punto sarebbe anche necessario modellare in modo più realistico il contatto perno/foro. Per completezza sono stati considerati anche altri parametri semplici, come la massima tensione circonferenziale, utilizzata in [13], che non tenendo conto della biassialità nello stato di tensioni non riesce a correlare i risultati, fig. 9. Risultati ancora peggiori si sono ottenuti utilizzando il parametro Crossland, già usato in [7, 9]. L'effetto della compressione trasversale sulla tensione idrostatica è così forte che si sono ottenuti anche valori negativi del parametro. Tornando ai risultati numerici, la tensione trasversale di compressione è danneggiante dal punto di vista della resistenza a fatica. Per meglio evidenziare questo aspetto sono state effettuate alcune prove utilizzando un perno avente una spianatura laterale, come mostrato nell'inserto di fig. 10; in questo modo si evita il contatto con il foro fino ad un angolo superiore all'angolo nel quale mediamente si erano manifestate le rotture utilizzando un perno circolare (39° nel caso i=1.25%). Questo accorgimento è stato utilizzato per evidenziare il problema del fretting nei collegamenti lug/fork, [14]. I risultati di queste prove, riportati nella fig. 10, mostrano il forte aumento di vita a fatica che si può ottenere con questo accorgimento. In questo caso il punto critico a fatica torna ad essere la radice dell'intaglio. Ulteriori prove, di sola propagazione, sono state effettuate su provini aventi una larghezza di 80 mm con lo scopo di evidenziare l'effetto dell'interferenza sulla propagazione dei difetti. Prima dell'installazione del perno con interferenza, è stato creato un difetto passante, simmetrico, avente una dimensione iniziale di 0.3 mm. La fig. 11 mostra i risultati di una di queste prove, effettuata con una tensione massima S=130 MPa, R=0.1. Per le stesse condizioni di prova, in assenza del difetto iniziale, si deduce dalla fig. 3 una vita di 2.5*106 cicli. Alla prima osservazione il difetto aveva una dimensione di 0.4 mm; poichè la rottura finale è avvenuta dopo circa 60000 cicli, si può concludere che l'interferenza è molto efficace nel ritardare la nucleazione dei difetti, mentre il suo effetto sulla propagazione è estremamente limitato. A conferma di ciò, si è simulata con AFGROW [15] la propagazione dello stesso difetto da un foro aperto, quindi in assenza di perno, ottenendo un risultato poco diverso, come mostrato in fig. 11 (curva a tratto continuo, senza simbolo). Attualmente non è disponibile una soluzione per il fattore di intensità degli sforzi per una cricca che si nuclea in un foro pieno; è comunque possibile utilizzare i risultati dei calcoli effettuati per valutare, anche se in modo approssimato, la propagazione di un difetto da un foro pieno. La previsione si basa sull'ipotesi che il difetto non modifichi le autotensioni presenti; questa ipotesi è accettabile per difetti piccoli ed in questa condizione viene consumata una cospicua aliquota della vita a fatica. Con il metodo degli elementi finiti, in assenza di difetti, si è valutata per il caso in esame (i=2.5%, Smax=130 MPa, R=0.1) una tensione sul bordo del foro variabile fra 118.8 MPa e -13.07 MPa. Si può idealmente rimuovere il foro e considerare una cricca, avente dimensioni 2a+D, che si propaga in tale stato di sollecitazione. La simulazione effettuata con AFGROW ha fornito un risultato, sempre riportato in fig. 11, in buon accordo con i risultati sperimentali. Dal punto di vista pratico il problema è 77 risultato essere leggermente più complesso, poiché in tutte le prove di propagazione la cricca si è nucleata da un solo lato, ma è semplice tenere conto nella simulazione anche di questa evenienza.

3. PROVINO CON TRASFERIMENTO DI CARICO

3.1 Descrizione del provino

La seconda parte del lavoro ha riguardato il caso di un provino con un foro nel quale è installato un perno che trasferisce carico, fig. 12; questa geometria è ovviamente più interessante dal punto di vista applicativo. Anche in questo caso si sono utilizzate delle bussole FTI per non danneggiare il foro durante l'introduzione del perno con interferenza. Volutamente il perno non ha serraggio assiale, per cui il carico viene interamente trasferito per taglio, massimizzando in questo modo la criticità del foro.

3.2 Risultati delle prove e loro elaborazione

I risultati delle prove, effettuate per tre diversi valori dell'interferenza, 0, 1.25 e 2.5%, sono riportati nella fig. 13, unitamente, come termine di confronto, ai risultati delle prove effettuate sul materiale base e sui provini con foro aperto; anche per questo provino si osserva un evidente effetto benefico dell'interferenza sulla vita a fatica. E' altresì evidente l'effetto negativo del trasferimento di carico in un foro. E' noto che una concentrazione di tensione è tanto più gravosa quanto più il carico è applicato vicino ad essa; il foro che trasferisce carico massimizza questa criticità. E' comunque interessante rilevare che alla maggiore criticità del foro caricato corrisponde un maggior effetto benefico dell'interferenza; ad esempio a 106 cicli un'interferenza i=1.25% produce un aumento del 22% della tensione ammissibile per il caso di foro senza trasferimento di carico, mentre tale valore sale all'82% nel caso di foro che trasferisce carico. Anche per questo provino si è effettuata la modellizzazione con il metodo degli elementi finiti per valutare lo stato di tensione prodotto dall'introduzione di perni con diversa interferenza e dalla successiva applicazione dei carichi di fatica. Come per il caso di foro senza trasferimento di carico si è valutata la tensione equivalente monoassiale, con la quale sono stati interpretati i risultati delle prove di fatica. I risultati ottenuti sono mostrati in fig. 14, confrontati con i risultati delle prove di fatica effettuate sul metallo base. Come nel caso di provino con rivetto che non trasmette carico, la tensione equivalente praticamente riduce ad una sola curva i risultati ottenuti per i due valori dell'interferenza. Questa curva differisce marcatamente da quella determinata in fig. 8, relativa al provino senza trasferimento di carico; il metodo non appare quindi generalizzabile. Anche per il secondo provino, in assenza di interferenza, e quindi con uno stato di tensione monoassiale, i risultati delle prove di fatica esaminati in termini di parametri locali, sono ragionevolmente deducibili dalla curva di fatica del metallo base. La fig. 15 mostra il risultato delle analisi al microscopio ottico delle superfici di frattura. Per ogni prova è stato misurato l'angolo, rispetto alla mezzeria del provino, al quale si è manifestata la rottura per fatica. Questi risultati sono forniti per i tre livelli di interferenza, unitamente al valor medio di ciascun gruppo. Anche se è presente una elevata dispersione, si osserva un aumento dell'angolo al quale si manifestano le rotture per fatica con il livello di interferenza. Come già accennato precedentemente, volendo migliorare il livello delle previsioni numeriche, sarebbe necessario anche ricercare intorno al foro la locazione in cui il parametro esaminato diviene più critico. Ciò complica notevolmente l'analisi dei risultati.

CONCLUSIONI

Sono state effettuate le prove di fatica su due diversi tipi di provini in lega leggera 2024-T351; comune ai due provini è la presenza di un foro in cui è installato un rivetto con interferenza. I due provini si differenziano per il carico trasferito dal rivetto, pari a 0% e 100% del carico applicato. Le prove di fatica hanno mostrato, ancora una volta, l'effetto benefico dell'interferenza sulla resistenza a fatica. Tale effetto è maggiore nel caso di foro che trasferisce carico. I modelli agli elementi finiti dei provini hanno mostrato che l'interferenza produce uno stato di tensione biassiale, caratterizzato da una tensione circonferenziale fortemente variabile nel ciclo di fatica e da una tensione radiale quasi costante. L'effetto della pressione radiale è quello di produrre una deformazione circonferenziale legata, tramite il modulo di Poisson, alla deformazione radiale. Si è immaginato cha tale deformazione circonferenziale fosse prodotta da una tensione circonferenziale fittizia, il cui valore è uguale al prodotto della tensione radiale per il modulo di Poisson. Questa tensione, sommata alla tensione circonferenziale, ha fornito una tensione equivalente monoassiale, con la quale si è valutata la resistenza a fatica. I risultati delle prove di fatica, interpretati con la tensione equivalente, vengono praticamente ridotti ad una sola curva, ma tale curva non 78

è la stessa per i due diversi provini. Il metodo non appare quindi generalizzabile. Inoltre tali curve non sono deducibili dalla curva di fatica del materiale base, mentre nel caso di interferenza nulla, per entrambi i provini, i risultati sono ragionevolmente deducibili dalla curva di fatica del materiale.

RINGRAZIAMENTI

Gli autori ringraziano lo studente Francesco Fruci per l'aiuto fornito nella esecuzione delle prove sperimentali.

BIBLIOGRAFIA

1. J. H. Jr Crews "An elastic analysis of stresses in a uniaxial loaded sheet containing an interference-fit bolt". NASA TN D-6955; 1972. 2. J. P. Hou, D. A. Hills "Interference contact between a pin and plate with a hole". Journal of Strain Analysis for Engineering Design, vol. 36, n. 5, 2001, pp. 499-506. 3. A. Lanciotti, C. Polese "The effect of interference fit fasteners on the fatigue life of central hole specimens". Lavoro accettato per la pubblicazione sulla rivista: Fatigue and Fracture of Engineering Materials and Structures. 4. ABAQUS Version 6.4, Copyrigth © Hibbitt, Karlsson & Sorensen, Italy. 5. C. G. C. Poussard, M. J. Pavier, D. J. Smith "Prediction of residual stresses in cold-worked fastener holes using the finite element method". American Society of Mechanical Engineers, Engineering Systems Design and Analysis, vol. 64, n. 8-1, 1994, pp. 47-53. 6. S. Larouche, M. Bernard, T. Bui-Quoc, D. Julien "Influence of cold working and interference fit on fatigue life of 7475-T7351 aluminium alloy fastener hole". International Committee on Aeronautical Fatigue, 21st Symposium, Toulouse, 27-29 June 2001, pp. 681-698. 7. D. Duprat, D. Campassens, M. Balzano, R. Boudet "Fatigue life prediction of interference fit fastener and cold worked holes". Intern. Journal of Fatigue, vol. 18, n. 8, 1996, pp. 515-521. 8. B. Crossland "Effect of a large hydrostatic pressure on the torsional fatigue strength of an alloy steel". Proc. Int. Conference on the Fatigue of Metals. The Institution of Mechanical Engineers, London, 1956, pp.138-149. 9. G. Shatil, A. G. Page "Multiaxial fatigue analysis of interference-fit steel fasteners in aluminum Al 2024- T3 specimens". American Society Testing Materials; Special Technical Publication, n. 1417, 2002, pp. 192-208. 10. D. L. McDiarmid "Shear stress based critical-plane criterion of multiaxial fatigue failure for design and life prediction". Fatigue and Fracture of Engineering Materials & Structures, vol. 17, n. 12, 1994, pp. 1475-1484. 11. Anon. "Metallic materials and elements for aerospace vehicle structures". Military handbook, MIL- HDBK-5H, 1998. 12. J. H. Jr Crews "Analytical and experimental investigation of fatigue in a sheet specimen with an interference bolt". NASA TN D-7926.1975. 13. J. H. Jr Crews "An elastoplastic analysis of a uniaxially loaded sheet with an interference-fit bolt". NASA TN D-7748. 1974. 14. J. Schijve "Fatigue of structures and materials". Kluwer Academic Publishers. 2001. 15. J. A. Harter "AFGROW user guide and technical manual". Air Force Research Laboratory, VA-WP- TR-2003-XXXX. 2003.

79

R =200 t=1.6 mm =4.00

270 20

Figura 2 – Dettaglio di un provino con perno montato all’interno di una bussola FTI.

Figura 1 – Provino senza trasferimento di carico.

350

300 Materiale base 250 i=1.25% Î Î Î i=0.75% i=1.875% 200 i=2.5% i=0.5% 150

Î Î Foro aperto

Smax [MPa] 100 Materiale base Foro aperto i=0.5% i=0.75% i=1.25% i=1.875% i=2.5% R=0.1 50 105 106 107 Cicli Figura 3 – Provino senza trasferimento di carico. Risultati delle prove di fatica. 80

500 0.05

i=0.5% l/l)

400 Sy ∆ 0.04 (

Sx ε 300 ey 0.03 ex 200 Von Mises 0.02

100 0.01

0 0.0 S (MPa) Figura 4 – Tensioni e -100 -0.01 deformazioni in prossimità di un foro in cui è installato un -200 -0.02 perno con bassa interferenza -300 -0.03 (i=0.5%). -400 -0.04 02468 X (mm) 500 0.05 l/l) 400 ∆ 0.04 i=1.25% (

Sy ε Sx 300 0.03 ey 200 ex 0.02 Von Mises Figura 5 – Tensioni e deformazioni in prossimità di 100 0.01 un foro in cui è installato un 0 0.0 perno con alta interferenza

S (MPa) (i=1.25%). -100 -0.01

-200 -0.02

-300 -0.03

-400 -0.04 02468 X (mm)

400 i=1.25% Interferenza 300 Interferenza + 150 MPa Interferenza + 15 MPa

200 S (MPa) y S y 100 x

Figura 6 – Tensione longitudinale durante le fasi di carico e scarico di un provino in cui è installato un 0 0246 perno con interferenza. X (mm) 81

300 S Figura 7 – Tensioni massime 200 agenti sul bordo del foro y conseguenti alla installazione 100 S di un perno con interferenza y x e alle successive fasi di carico 0 e scarico del provino. -100 S=0

S (MPa) -200 S=170 MPa S=17 MPa -300

-400 S x -500

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 interferenza (%) 350 FORO APERTO 300 MATERIALE BASE

(MPa) 250 i=0.5% i=0.75% Î Î Î R=0.1 i=1.25%

Seq 200 Î Î Foro aperto i=0.75% Figura 8 – Provino senza 150 i=1.25% i=1.875% i=2.5% trasferimento di carico. i=2.5% Risultati delle prove di fatica Materiale base in funzione della tensione i=0.5% equivalente. i=1.875% 100 105 106 107 Cicli

350 Foro aperto 300 i=0.5% Materiale base 250 i=0.75% Î Î Î

Figura 9 – Provino senza 200 trasferimento di carico. i=1.25% Risultati delle prove di fatica Î in funzione della tensione 150 circonferenziale massima. i=1.875% Î Tensione circonferenziale massima (MPa) massima circonferenziale Tensione i=2.5% 100 105 106 107 Cicli 82

200 i=1.25%; R=0.1

Figura 10 – Provino senza trasferimento di carico. Risultati delle prove di fatica con perno spianato. Perno circolare 150 Perno spianato Smax (MPa) 4.05

3.65

Direzione del del carico Direzione Perno spianato 100 105 106 107 Cicli Smax=130 MPa, R=0.1 12

a0 D a0 10 a (mm)

8 as ad Figura 11 – Propagazione di un 6 a=(as+ad)/2 difetto da un foro con installato un perno con interferenza; risultati sperimentali e previsioni numeriche. 4

2 Foro pieno, i=2.5%. (Sperimentale) Foro vuoto. (Previsione AFGROW) Foro pieno; i=2.5%. (Previsione: Smax=118.8 MPa, R=-0.11) 0 0 10000 20000 30000 40000 50000 60000 Cicli Afferraggio 4

20 Provino 1.6

Figura 12 – Provino con trasferimento di carico. 83

300

Figura 13 – Provino con trasferi- MATERIALE BASE -> -> -> mento di carico. Risultati delle 200 prove di fatica per diversi valori FORO APERTO dell’interferenza fra foro e perno.

100 90 80 70 60 i = 1.25% Smax (MPa) 50 i = 2.5%

40 i = 0% 30

20 104 105 106 107 Cicli 350 325 300 MATERIALE BASE 275 250 -> -> -> 225 i = 0% 200 Figura 14 – Provino con 175 trasferimento di carico. Risultati delle prove di fatica 150

(MPa) i = 2.5% in funzione della tensione 125 equivalente. i = 1.25% eq R=0.1 S 100

75

104 105 106 107 Cicli

30

20 Figura 15 – Posizione della Cricca sezione critica in funzione dell’interferenza. 10 (°) θ

0

-10

0.0 0.5 1.0 1.5 2.0 2.5 Interferenza (%) 84

Numerical Simulations and Experimental Evaluations of some Technological Processes Employed to Improve the Fatigue Resistance of Stressed Holes

Luisa Boni*, Francesco Pietra**, Claudia Polese* * Department of Aerospace Engineering, University of Pisa, Pisa, Italy ** EnginSoft S.p.A, Florence, Italy E-mail: [email protected] Web: http://dottorato.dia.ing.unipi.it/Dottorandi

Summary Different technological processes had been developed to enhance the fatigue resistance of critical holes: they are mainly based on the introduction of a residual stress field around the hole that can reduce mean stress or stress amplitude of the applied fatigue spectrum. Aeronautical industries accept as a standard practice Split Sleeve Cold ExpansionTM and Interference-Fit methods for riveted joints and ForceMate® bushing installation for open holes or structural connections as lug-fork. The investigated techniques are obviously different from the technological point of view but quite similar in the theoretical and FEM analysis aspects. In order to provide an effective design tool, different FEM models of these three techniques, characterized by increasing degrees of approximation of the real processes, had been developed. The obtained stress results are really helpful to investigate process parameters influence in enhancing fatigue life and to formulate some simple theoretical provisional methods. The evaluated fatigue life is finally compared with series of experimental tests data on typical aeronautical material.

1 85

Keywords Interference-Fit, Split Sleeve Cold ExpansionTM, ForceMate®, residual stresses, FEM analysis, theoretical methods, experimental results.

Introduction The fatigue resistance of aircraft structures is directly connected with the fatigue resistance of structural joints. Focused only on mechanical joints a possible distinction can be made between permanent connections, which are usually realized by means of solid rivets, and threatened fastener or mechanical joints that can be dismounted. In order to enhance fatigue life of designed high stressed or demonstrated in service critical holes different techniques had been developed. Among others three similar processes had been analyzed: the shared characteristic is the initial introduction of an oversized pin into the processed hole. In Interference-Fit process the pin (or rivet) is permanently installed; the interference level can be chosen in order to produce only elastic strains surround the hole. In this way the resulting elastic stresses field reduces the fatigue cycle stress amplitude. Instead in Split Sleeve Cold ExpansionTM process the interference level is high enough to produce plastic strains and the pin is subsequently extracted. So the residual stress field around the hole decreases the fatigue cycle mean stress. In ForceMate® process both the positive effects are included: the installation of an high interference bush produces also the hole cold working. Besides advantages associated to the compressive residual stresses field, the interference mounted bush avoids the direct contact between the hole and possible connection pin. A large research activity started at the Department of Aerospace Engineering of the University of Pisa dealing with the experimental evaluation and numerical simulation of these processes in the framework of the European project ADMIRE and in cooperation with some italian aeronautical industries.

Experimental Activity and FEM Models During the experimental activity an increased fatigue life had been observed for higher squeezed riveted joints, with a D/d ratio between the diameter D of the formed head and the initial diameter d of the rivet shank higher than 1.2, i.e. 1.4 and 1.6 [1]. Actually when a solid rivet is driven the second head is formed but the rivet shank is also plastically deformed so the hole is filled creating a beneficial interference state. In order to evaluate the stress field in the joint a complete analysis is possible with a FEM process simulation. The analysis immediately appears quite complex due to many parameters to taking into account during the simulations. On the other hands few check parameters are disposable for evaluate the simulation quality. Is possible to measure the dimension of the former head or the strength or the position of the driven head machine during the process. Despite that is quite difficult to introduce the correct elastic plastic behavior of the material and quite big differences in the solution are obtained if the remeshing capability is or not introduced. To clearly understood the phenomenon and elaborate a good design instrument the simple case of an unloaded fastener filled hole, commonly named “zero load transfer” was examined. A lot of this joints exists in aircraft structures such as connection between longitudinal stiffeners and external fuselage skin, so this

2 86 simple case is of practical interest too. Tests had been performed realizing a simple strip specimen with a patch in which rivets had been installed with the same previous squeeze values. In order to avoid possible load transfer due to friction a Teflon layer was inserted between the two sheets. The interference value had been so experimentally evaluated from different sections of the rivet steam. The interference value measured nearest to the observed critical area had been inserted in simplified FEM models with the “clearance/gap” command. As expected results shown a reduced stress range in the critical area during the load application compared with the no interference rivet. So the enhanced fatigue life is highly connected with the reduced stress amplitude during the test. Unfortunately obtained stresses introduced in analytical program didn’t give rise to good previsions of the fatigue life experimentally observed. The problem seams to be still quite complicate: in the real situation the rivet stem is highly deformed so the contact is not constant around the hole, the interference level is not the same along the stem and is not exactly the same in all riveted holes. In order to simplify the problem several tests had been carried out on dog bone specimens having a central hole filled with different pins in order to obtain 5 different interference levels in the range from 0.5 to 2.5%. The steal pin was the same for each interference levels in order to avoid possible scatter in the results connected to the tolerance on pin machining. The hole was protected from possible damage during the pin insertion by a self-lubricated split sleeve so no wear was observed in the pins. Two different FEM models had been elaborated. The most complex simulates the real insertion of the pin and the other simulates only the interference by means of the clearance command. Stresses obtained in the critical diametrical area along a radial path show practically the same results. So in the FEM analyses the simplest model was preferred being less time consuming. Circumferential strains obviously increase with higher interference values, instead circumferential stress after an 1% interference level starts to decrease due to material plasticity. An experimental verification of the obtained results is quite complex according to the small specimen dimensions. A biggest specimen was instrumented by 16 strain gauges positioned near the hole along and across the load direction. A biggest pin was interference mounted by means of an hydraulic jack and the specimen was loaded. Experimentally measured strains shown a good agreement with FEM results in which the experimentally measured stress-strain curve of the base material was inserted. The same calculation was performed for fatigue tests specimen. Seven interference levels were examined from 0 to 4 % and different applied stresses covering the range of stresses applied in fatigue tests. In each model three different steps were examined: pin insertion, application of the maximum stress and unloading to the minimum stress, R=0.1. Interference Fit produces a biaxial stress state composed of a longitudinal stress, whose variation constitutes the actual fatiguing stress and a transverse stress, whose value, at a fixed interference degree, is quite constant in fatigue cycles. Moreover, after the pin insertion and the first fatigue cycle, subsequent fatigue cycles, at the same level, only cause elastic stresses. So hoop strain seams to be the simpler and representative parameter that can be converted into an equivalent uniaxial stress, whose minimum and maximum values in the fatigue cycle are given by Seq min = Symin − νSxmin and Seq max = Symax − νSxmax. The direct comparison of the fatigue results can be made only correcting the effective stress ratio of each tests using the relationship given in [2] to evaluate the equivalent maximum stress in the case R = 0.1. The method correlates very well the fatigue test results obtained under different interference levels, which are almost reduced to only one fatigue curve. But some additional factors are present, such as fretting and the position of the critical locations: the examination of the fractured specimens shows, besides a large scatter in the results, that the critical location

3 87 moves around the hole by increasing the interference level up to 40–50°. So, for any criterion adopted, calculations would have to be carried out in several locations around the hole, looking for the critical value of the parameter. Some tests were carried out with faceted pins in order to realize a uniaxial stress state at the notch root: results demonstrated that the suppression of the transverse stress increases fatigue life and what's more the fatigue critical location is again the notch root. The equivalent uniaxial stress evaluated by FEM models is able to correlate experimental results quite well, but placed in slightly over the open hole fatigue results. Moving toward the simulation of more realistic problems also a load transfer specimen had been experimentally tested for three different interference levels. The specimen was realized by an interference pin insertion with an FTI sleeve interposition as previously done. The pin was not tightened in order to avoid load transfer by friction. In this way the notch root is in the more critical condition. The experimental results put in evidence the beneficial effects of interference fit: an interference percentage of 1.25% increase the design stress at 106 cycles life of 22% for a no load transfer joint and of the 82 % for a load transfer joint. Also for this test a FEM model had been realized in order to evaluate stresses introduced by pin insertion and subsequently loading and unloading. Results interpreted in terms of equivalent stress practically became a single line, but cannot be directly derived from the base material curve. Fracture surfaces observed by means of microscope clearly shown as previously an increased angle of preference fracture at higher interference levels. This point so needs to be deeply investigated. Other systems for enhance fatigue life developed by Boeing Company and distributed by Fatigue Technology Inc. (FTI) are now accepted as a standard practice in the aeronautic industry. Split Sleeve Cold ExpansionTM is accomplished by pulling a tapered mandrel, pre- fitted with a lubricated split sleeve, through a hole. The function of the disposable split sleeve is to reduce mandrel pull force, ensure correct radial expansion of the hole and preclude damage to the hole. The action of drawing the mandrel through the starting hole causes a radial plastic flow of material and results in an annular zone of beneficial residual compressive stresses that extend up to one diameter beyond the edge of the hole. At the end the hole is reaming to final size. A wide experimental activity had been carried out at the Department of Aerospace Engineering to evaluate the influence of the cold working process on fatigue resistance of riveted joints. The problem of the quantitative assessment of compressive residual stresses around the hole has been faced by several FEM analyses characterized by different degrees of approximation in the process simulation. The three dimensional nature of the cold expansion process has been accounted for in all analyses, by considering 3-D models. Due to the highly strains on the material surrounding the hole an experimental evaluation of the complete stress- strain curve of the employed material was obtained. The considered material agrees fairly well with the kinematic plastic hardening model, although a remarkable “Baushinger’s effect” in compression. The FTI specifications for the cold expansion process can be divided into three main steps: hole expansion by pulling a tapered mandrel through the hole, hole recovery once the mandrel is removed and a finishing ream that removes a thin layer of material around the hole. The differences among the considered analyses are relevant to the first two phases, while the third phase is in any case simulated by the command “element removal/kill”. The hole expansion can simply be simulated by imposing and removing an uniform radial displacement on the nodes at the hole edge or the mandrel is directly modeling and its drawing across the hole. The presence of a lubricated split sleeve can be accounted for o not. The FEM distribution of residual tangential stresses agree fairly well with those obtained by using the experimental drilling method proposed by Sachs, [3]. The more complex models can put in evidence some side effects of the process no present in the simplest ones. The introduction of the lubricated

4 88 split sleeve, that mediates the contact between hole and mandrel causes a loss of axial- symmetry. Besides the simulation of the actual mandrel drawing introduces a variation of the residual stress field in the plate thickness: the tangential component is minimum at the entry face, afterwards it increases up to the absolute maximum value at about the middle distance, keeping higher at the exit face. The effect of a double application of the cold expansion process in opposite directions has been also investigated observing an improvement, in terms of tangential component, characterizes about the four fifth of the plate thickness. Fatigue and propagation tests had been performed on “dog-bone” specimens with a central hole no cold worked, simple and double cold worked. For cold worked specimens the two states of initial defect lying on the entry (first drawing) face and on the exit face have been considered. Experimental results, consistent with FEM analyses, put in evidence the strong retardation effect of cold working on crack propagation. Cold worked specimens life is, respectively, 10 and 30 times greater than life of the pristine specimens, respectively in the cases of initial defect at the entry and exit face. In the case of double cold working, life increases about 60 times for the entry face and more than 100 times for the exit face. A big work had also been done for the characterization of the ForceMate® FTI process on the AerMacchi M-346 training fighter. The connection between the main fuselage frame and the wing spar is realized by a lug-fork connection. The severe fatigue cycle spectrum to which the component is subjected required a good design and a damage tolerance approach. ForceMate provides an alternative to shrink or press fit bushing installation methods with significant life improvement. The bushing is initially placed in the hole as clearance fit, and then using an oversized tapered mandrel, the lubricated ForceMate bushing is cold expanded into the hole. The FEM simulation is conducted for the entire process in order to evaluate the effectiveness of the process itself. Obtained strains are in good agreement with strain gauges measurements achieved during bush installation. The system seems to be really effective also in introducing a cold working of the work piece. Also tests conducted with the real fatigue spectrum shown the highly increased fatigue life of the processed specimen compared to the one without ForceMate. The cold working of the hole is clearly demonstrated also in one test in which the ForceMate bush had been mechanically removed. The improved fatigue life is more than two times also in presence of an artificial flaw quite big.

FEM simulation: insertion of an oversized pin

5 89

Conclusions

Some FEM analyses, described in the present paper, have been performed to simulate Interference Fit, Split Sleeve Cold ExpansionTM and ForceMate® processes with the aim of providing useful instruments for the estimation of the service life enhancement of stressed aerospace components. Good correlations of the numerical strains with experimental strain gauges verification had been obtained, and the quantitative evaluation of the enhanced fatigue life, despite some complexity, had been experimentally demonstrated. A tuning of the FEM model is necessary, specially for a correct material model definition. Simplest models can be really effectives, but more time consuming models can highlight some important side effects and clarify experimental results. All the analyzed method shown to be successful in enhancing fatigue life in simple specimen. More detailed analyses of the real geometries can be of practical interest.

References

[1] COMPETITIVE AND SUSTAINABLE GROWTH Programme, Advanced Design concepts and Maintenance by Integrated Risk Evaluation for aerostructures, acronym ADMIRE, Project N°GRD1-2000-25069, Contract N°G4RD-CT-2000-00396: G. Cavallini, C. Polese “Final Report on Tests Carried Out for the Evaluation of Short Crack Effect”. Document N°. ADMIRE-TR-3.0-61-3.1/UP., New Perspectives in Aeronautics, Submission date 31/07/03.

[2] Anon “Metallic materials and elements for aerospace vehicle structures”, Military Handbook, MIL-HDBK-5H, 1998.

[3] Sachs G., Van Horn K. R. “Practical Metallurgy: Applied Physical Metallurgy and the Industrial Processing of Ferrous and Nonferrous Metals and Alloys”, American Society for Metals, Cleveland OH, 1940.

6 90 Chaboche Material Model Optimization for Evaluation of Residual Stress Field in Cold Worked Holes

Claudia Polese, Sergio Urso Sculco

Department of Aerospace Engineering, University of Pisa, Pisa, Italy

Francesco Franchini, Francesco Pietra

EnginSoft S.p.A., Firenze, Italy

Summary

The present paper describes work undertaken to correctly evaluate fatigue life enhancement of Split Sleeve Cold worked holes in an aerospace aluminum alloy. The “Split Sleeve Cold ExpansionTM” process is a widespread cold working technique capable of generate a compressive residual stress field around treated holes, thus potentially providing an increased cost efficiency and durability in metallic aero-structures. However, an experimental and predictive knowledge of fatigue behavior of cold worked holes must be firstly obtained if the technique is to be included in damage tolerance of structural design and optimization, as well as in service maintenance. A teamwork of the Department of Aerospace Engineering (DIA) at the University of Pisa and EnginSoft S.p.A. Florence carried out both experimental and numerical activities relevant, as a first step, to the simple case of open hole without load transfer. Fatigue and crack propagation tests, performed on dog bone specimens in 7075-T73 aluminum alloy, thickness 2.3 mm, having a central hole of 4.83 mm diameter, demonstrated a remarkable increment of the fatigue life due to cold working process. The problem of the quantitative assessment of the complex three-dimensional compressive residual stresses field was faced by means of strongly non-linear Ansys Finite Element Analyses. Particularly more attention was focused on modeling the three dimensional nature of the cold expansion process and the real plastic material behavior. The “Split Sleeve Cold ExpansionTM” was simulated by modeling a complete mandrel drawing across the hole with an interposed sleeve. In order to obtain a realistic description of the stress fields during process phases was necessary to introduce a Chaboche plastic material model. In fact, a general kinematic hardening material behavior was not able to describe the strong Bauschinger effect exhibited by the aeronautical aluminum alloy during the unloading phase, giving rise to a non conservative compressive residual stress field around the hole, inconsistent with experimental fatigue test results. A valuable Ansys Chaboche material model was implemented joining the structural implicit code Ansys with the multi-disciplinary and multi-objective design optimization tool, modeFRONTIER. The model was obtained by superimposition of five kinematic hardening models, tuned with experimental static tests on the base material subjected to a stress-strain history similar to the material at the cold worked hole edge. The numerical results of the complete 3D FEM analysis highlighted asymmetric three dimensional compressive residual stresses along the circumferential and through thickness directions at the hole, confirmed by experimental data of defect nucleation and propagation. A realistic description of the stress fields during process phases was so achieved and, at the same time, a useful instrument for similar applications was developed.

Keywords: Cold working, Split Sleeve Cold ExpansionTM, residual stresses, Ansys FEM analysis, kinematic hardening, Baushinger effect, Chaboche material model, modeFrontier optimization, experimental results.

modeFRONTIER Users’ Meeting 2006

September 28-29 2006, Trieste – Italy

Aerospace, Automotive and Marine Applications 1 91

0. Introduction

In aerospace industry fastener holes are considered as potential crack initiation sites for structures that undergo significant cyclic fatigue loads because of the high stress and strain concentrations. Cold working of critical hole locations is now widely practiced to increase the service life of metallic aero- structures. The main idea behind the cold working process is to introduce a compressive residual stress zone in the material around the hole. This compressive stress field improves fatigue performance by retarding the crack propagation rates of flaws that originate from the hole. The Split Sleeve Cold ExpansionTM technique, developed at the Boeing Company and marketed by Fatigue Technology Inc. (FTI), has been accepted as a standard practice in the aeronautic industry. Go over the main points the process is accomplished by pulling a tapered mandrel, pre-fitted with a lubricated split sleeve, through the hole, Fig.1. The combined thickness of the sleeve and the mandrel major diameter radially expand the hole. The lubricated sleeve shields the hole from frictional forces generated by the high interference of the expansion mandrel in order to reduce the pull force required, protect the hole, and allow the mandrel to exert radial forces through the entire thickness of the material field, so resulting in the desired annular zone of residual compressive stresses. At the end the hole is reamed to final size. The magnitude of the peak residual compressive circumferential stress is usually about equal to the compressive yield stress for the material. The compressive stress zone spans one radius to one diameter from the edge of the hole, for most materials. A balancing zone of tensile stresses lies beyond the circumferential compressive stress zone. The present study of cold worked holes is motivated by several practical considerations. First, for both future structural design of aerospace structures as well as their inspection and repair schedules it is a paramount importance to be able to predict the fatigue life of cold worked holes in aero-components reliably. The main difficulties are the quantitative assessment of the existing residual stresses introduced by the Split Sleeve Cold ExpansionTM process due to the three dimensional complex nature of the expansion process itself and the correct definition of the material elastic-plastic behavior. In order to improve the process knowledge a focused experimental activity was performed at the Department of Aerospace Engineering at University of Pisa in the framework of a cooperation with EnginSoft S.p.A. Florence to develop a complete numerical research program.

1.0 Experimental activity

In order to evaluate the effect of the cold working process the aeronautical aluminum alloy 7075-T73 was selected, typical of primary structures at high stress levels. Some plates of thickness 13 mm and 2.3 mm were used during the experimental activity. In joints of two plates of thickness 2.3 mm a rivet or bolt of 4.83 mm diameter is a usual application. So from the “FTI tooling catalog” of the cold expansion systems the starting hole and tool diameters were sized to give a finish hole diameter compatible with clearance fits of standard fastener. For a finished reamed hole of diameter size 3/16” a 4-4-N mandrel was selected with a starting hole of 4.3 mm in order to obtain an interference level between the mandrel and the treated hole of about 4%. As a first step toward the problem solution an experimental evaluation of the complete tensile- compressive stress-strain curve of the material, aluminum alloy 7075-T73, was obtained. Correct data are strictly necessary because of the highly positive-negative strains recorded in the material surrounding the hole during the cold working process. Material data were obtained by means of static tests on specimens with a circular cross-section of 6.35 mm diameter, Fig. 2a, and related to the stress-strain curve shown in Fig. 3. The basic behavior under only tension or compression is compared with the results of more detailed tests. Some specimens were stressed in tension until a fixed value of total deformation (i.e. about 2, 3 and 4%) and, successively, they were stressed in compression. The stress-strain history imposed to the specimens was selected in order to replicate the stress strain history of different points near the cold worked hole edge. As a matter of fact, the yielding in tension occurs quite abruptly, while the compression branch exhibits a significantly more gradual behavior. The remarkable decrease of the yielding stress in compression following a yielding in tension attests a strong ‘Bauschinger’s effect’ of the material, i.e. a strong diminishing of the compressive yield stress and a smoothing of the negative part of curve. In the same graph of Fig. 3, the curves relevant to the most common models for the material plasticity are also reported, that are the bi-linear kinematic and isotropic hardening models. It is evident that the considered material agrees fairly well with the kinematic hardening model, although remarkable differences are present

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Aerospace, Automotive and Marine Applications 2 92 between the behavior in tension and compression. An improvement on material models seems to be strictly necessary in order to correctly evaluate the residual stress field around the hole. In order to obtain a quantitative assessment of the influence of the selected process, fatigue tests on simple cold worked open hole specimens, without load transfer, were conducted. In particular, fatigue and propagation tests were performed on ‘dogbone’ specimens with a central hole, made of 7075-T73 aluminum alloy, with 2.3 mm thickness and fibers in the longitudinal direction, Fig. 2b. The sleeve was inserted in the initial hole, diameter 4.3 mm, with the split in the load direction. The cold working was performed by means of a FTI 4-4-N mandrel, and the final reaming at the 4.83 mm diameter. The lowest value of reaming was selected from the user manual in order to achieve the highest residual stress field around the hole. Fatigue tests were performed on pristine specimens and specimens subjected to simple and double cold working, performed in the same or opposite directions. In propagation tests an initial corner flaw was introduced, starting from which fatigue cracks nucleated. For the cold worked specimens, the two cases were considered of initial defect lying on the entry and the exit faces. As far as the specimens subjected to double cold working are concerned, the entry face is the one relevant to the first mandrel drawing. For all the tests a servo-controlled machine was used having a 10 KN load capacity. For the propagation tests only, the surface size of the defects was monitored by means of an optic microscope 40x, mounted on a rail and equipped with a position visual display unit having a 0.01 mm resolution. In Fig. 4 are reported result relevant to the fatigue test, performed at different maximum stress levels between 135 and 190 MPa with a stress ratio R=0.1. The Split Sleeve Cold ExpansionTM greatly enhances fatigue life of open holes respect to the pristine ones. A significant improvement of the number of cycles to final failure is obtained also applying the process two times in concord or opposite directions. Obviously in this case the process cost is almost twice than the single one, and some problems of accessibility to the structure must be considered. All crack propagation test were performed at the same stress level Smax=155 MPa, R=0.1. Results in Fig. 5 shown crack propagation data relevant to a same value of initial crack length. It is evident the strong retardation effect of the cold working on the crack propagation. In particular, the life of the cold worked specimens is, respectively, 10 and 30 times greater than the one of the pristine specimens, for the cases of initial defect on the entry and the exit face. This behavior confirms the different effectiveness of the cold working at the two sides of entry and exit of the mandrel. In the case of double cold working, the life increases of about 60 times for the entry face and of more than 100 times for the exit face. A second application of cold working in reverse direction is thus able to improve the fatigue resistance on both sides. As far as the double cold working performed in concord directions is concerned, for the case of crack on the exit side, a fatigue life slightly higher than the one relevant to simple cold working was observed; on the contrary, for the case of initial defect on the entry side of the mandrel, the crack, once it reached, with decreasing growth rate, a size of 0.6 mm, didn’t further propagate, until the end of the test at more than 2.5ּ106 cycles. Indeed experimental results confirm also the positive effect of double cold working performed in concord directions.

2.0 Optimization of the Chaboche material model

In order to obtain an improvement respect to previous works relative to the evaluation of the residual stress field in cold worked holes, the first step is to optimize the material model with the base material tests as a reference. The easiest correlation check is based on the comparison between numerical results and experimental static data in similar conditions. During cold working expansion a circumferential deformation of about 4% is imposed to the material near the hole. When the mandrel leaves the hole the elastic spring-back of the surrounding zone causes a re-yielding in compression of the edge material. In order to reproduce material behavior at the hole edge some very simple coupons, Fig. 2a, were static tested, Fig.3. Specimen number 3 was stressed until a total deformation of 4.2% and after unloaded and loaded in compression till a negative deformation of -1.4%. As can be observed from the strain-stress curve the yielding in tension is well defined at 465 MPa. On the other hand yielding in compression after high plastic strains exhibits a strong Bauschinger effect with a value less than half, 215 MPa, and a completely different slope of the plastic part of the curve. This strain history is the most similar to the one experienced in present cold worked hole. Coupon number 2 was deformed till a 3.2% in tension in order to have also a reference curve for the material at a certain distance from the hole edge. A simplified square bar with a length of 100 mm was modeled into the Ansys finite element code, Fig. 6, with 320 quadratic elements SOLID186. The numerical analysis reproduces boundary

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Aerospace, Automotive and Marine Applications 3 93 conditions similar to the experimental test. One square face was fixed and at the other was imposed a longitudinal displacement to follow the same strain history. A positive uz of 4.2 mm was imposed at the first step, after the boundary condition was inactivate, so unloading the model, and subsequently a negative displacement was imposed till -1.4 mm. The non linear option were obviously activated (NLGEOM=ON, 20 substep) in order to evaluate the effect of highly plasticity. The material of the bar was modeled like an elastic isotropic material, with a Young Module = 70200 MPa. To define the inelastic behavior a non linear kinematic hardening model was selected, the Chaboche model, Fig.7. This Ansys option (CHABOCHE) typically uses the Chaboche model for simulating the cyclic behavior of materials. But like the bilinear kinematic and the multilinear kinematic (BKIN and MKIN) options is possible to use this model to simulate monotonic hardening and the Bauschinger effect. In addition is possible to superpose up to five kinematic hardening models and an isotropic hardening model to simulate the complicated cyclic plastic behavior of materials, such as cyclic hardening or softening, and ratcheting or shakedown. In order to completely define a numerical model of Chaboche material the following constants must be defined:

Constant Meaning C0 k = Yield tension stress

C1 C1 = Material constant for first kinematic model

C2 γ1 = Material constant for first kinematic model

C3 C2 = Material constant for second kinematic model

C4 γ2 = Material constant for second kinematic model ......

C(2NPTS-1) CNPTS = Material constant for last kinematic model

C(2NPTS) γNPTS = Material constant for last kinematic model

k, and all C and γ values in the left column are material constants in the Chaboche model. The maximum number of constants, m is 11, which corresponds to 5 kinematic models [NPTS = 5 on the TB command]. The default value for m is 3, which corresponds to one kinematic model [NPTS = 1].

The total material model was so implemented in the Ansys code as follow:

…... MPTEMP,1,0 MPDATA,EX,1,,70200 MPDATA,PRXY,1,,.33 TB,CHAB,1,1,N_MODEL5, TBTEMP,0 TBDATA,,C0,C1,C2,C3,C4,C5 TBDATA,,C6,C7,C8,C9,C10, ……

The challenge was to find correct values of these constants in order to obtain a good agreement with experimental test. That represented an hard manual task: consequently the optimization of the Chaboche material model was obtained by modeFRONTIER 3.1.3 code. ModeFRONTIER is an integrated multi-objective and multi-disciplinary design optimization software platform, and it can provide invaluable help for carrying out the tasks outlined above.

ModeFRONTIER can handle in automatic way the variable parameter (i.e. material constants) changing them to reach the best solution for describing material plastic behavior. The only request is to interface the software with a parametric FE model. Once the optimization problem is defined

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Aerospace, Automotive and Marine Applications 4 94 through an intuitive graphical user interface, it is easy to interface modeFRONTIER with Ansys (as also the greatest part of commercial code) by means of ASCII file. The only constraint is that the package must be able to compute the outputs from the inputs without requiring human intervention. At this point modeFRONTIER finds optimal designs by using a combination of optimization technologies and strategies, which can be classified as follows:

• Preliminary Exploration Methods: used to provide an initial sampling of the design space, include algorithms derived from Design Of Experiments (DOE) theory.

• Multi-Objective Optimization Methods: used to find designs that lie anywhere in the design space but near the Pareto Frontier, include genetic algorithms and evolution strategies.

• Local Refinement Methods: used to investigate the space around a given design, these single objective methods look for a local maxima or minima employing so-called Hill Climbing algorithms.

The approach used by modeFRONTIER to find optimal designs can be controlled by the user and can vary from the use of a single technique, to an hybridization of several techniques (e.g., one or more DOE methods can be used to generate an initial population of designs, which is then evolved using a Genetic Algorithm, and whose best candidates are then refined through an Hill Climbing algorithm). The present calibration problem was structured as shown in the work flow of Fig. 8: 11 input material variables (1 constant and 10 variable) are inserted in the parametric Ansys input file of the square bar. From the FEM results two outputs, the positive and the negative parts of the stress-strain curve, are compared with the experimental ones in order to minimize the differences. For each input variable, the modeFRONTIER user has to define a suitable range and a step of variation from which the DOE node allows to specify the first set of design. For the material variables relative to the C odd parameters (C1, C3,…) a range between 50 and 20000 was chosen, with a step of variation of 50. The γ even variables (C2, C4,…) were defined in a range -20 – 300 with a step variation of 1, Fig. 8. The values are in a possible wide range for Chaboche applications with small increments in order to cover a large part of the solution space. The DOE of 36 configurations was generated using a factorial Latin Square algorithm, Fig. 9. The method is effective at processing an high number of variables and generates a number of design equal to the squared number of levels, fixed to 6. DOE designs were so processed by the external script node (e.i. ANSYS) that imposed to the square bar the strain history in tension and compression described above. The FEM results obtained in the node 1321, chosen in the middle of the square bar, were processed by an Ansys Macro (VOPER) in order to obtain separately the integral of the positive and the negative parts of the stress-strain curve described by the node during the different steps. These values were compared with the experimental data in an Excel file generating the square of the absolute value of the differences between the FEM Voper integrals and the experimental integrals of the stress-strain curve, chosen as the two objectives to be minimized, Fig. 9. On the base of the FEM results the genetic algorithm started the optimization process: the optimization algorithm chosen is MOGA (Multi-Objective Genetic Algorithm) generational evolution and the number of generations is set up at 10, Fig. 10. The scatter chart of the design space, Fig. 11, shows a well defined Pareto Frontier, the field of the optimized design solutions. On the Pareto Frontier 9 well spaced optimum designs were chosen and the relative FEM curves were compared with the experimental one. Among other the design 192 is the better compromise to satisfy our goals, Fig. 12. The unloading material behavior of the optimized design unquestionably better describes the 7075-T73 stress-strain curve if compared with the two classic material models previously presented, Fig. 3. The 10 variables proved a low degree of correlations, as shows in the correlation matrix, Fig. 13, thus each one had a tangible significance in the two objectives, Fig. 14. The first two kinematic materials parameters (C1, C2) seem to have a smaller influence on the final results if compared with the other constants, i.e. C7 and C10, but no one kinematic model has a negligible effect. So an high confidence level can be referred to this optimized material model solution.

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Aerospace, Automotive and Marine Applications 5 95 3. FEM evaluation of the residual stresses in the material surrounding a cold worked hole

In order to obtain a quantitative assessment of the effective residual stress field was performed an Ansys FEM analysis capable to take into account the three dimensional nature of the Split Sleeve Cold ExpansionTM technology. The FTI specifications for the cold expansion process can be effectively divided into three key steps: hole expansion by pulling a tapered mandrel through the hole, hole spring back once the mandrel is extracted and a finishing ream that remove a thin layer of material around the hole. In previous similar analyses present in literature the hole expansion and its upturn has been simply simulated by imposing and removing an uniform radial displacement on the nodes at the hole edge. Often also the presence of the split sleeve is omitted or just incorporated in the mandrel diameter. In this way the problem result completely symmetric in the thickness and in the circumferential directions and the effect of the driving mandrel are completely lost. This solutions can hardly justified the experimental results observed during crack propagation tests. In order to increase the process knowledge in the present work a complete FEM model was developed: the employed mandrel was directly modeled and it was drawn across the hole. The lubricated split sleeve was interposed between the mandrel and the hole as in the real treatment of the hole. The final reaming was obtained by the ANSYS command EKILL (element kill), in order to remove the inner layer of material. The compressive residual stress field was estimated in the material surrounding the hole of a circular plate with a 50 mm diameter and a 2.3 mm thickness. The diameter of the plate was selected to be quite large with respect to the hole diameter, being the ratio of the diameters equal to about 10; it had been observed that the residual stress field is not significantly different for the case of smaller plates, having, for instance, the diameter equal to the width of the fatigue specimens, i.e. 25 mm. The thickness of the plate was selected equal to the thickness of the specimens tested. The other dimensions involved in the model geometry followed exactly the FTI specifications and were experimentally verified on the employed FTI instrumentation, Fig. 15: the initial hole with a diameter of 4.3 mm is relevant to the FTI 4-4-N mandrel and to the split sleeve of thickness 0.152 mm. The final reaming of the hole brings the diameter to the final size of 4.83 mm. The mapped mesh of the plate was realized with brick 8 nodes elements (element SOLID185) sweeping a surface mesh with not-solved element 4 nodes (element MESH200). A spacing ratio of 0.2 was applied to obtain a finer mesh near the hole. A circular partition of 0.35 mm at the hole edge was introduced in order to remove the elements during the final reaming. A regular mapped mesh was obtained also in the others two components in order to avoid fluctuations in the results connected to a no uniform node distribution. In Figs. 16 a-b are shown the mesh relevant to the groups plate-mandrel and plate-sleeve. The boundary conditions on four lines in the border plate thickness allowed the radial expansion of the plate, avoiding rigid body movement in the x-y plane. According to the process FTI specifications, during the mandrel pulling, a nosecap with circular cross-section, connected to the puller unit, is held firmly against the work piece. As a consequence on a circumference on the plate exit face, coaxial to the hole, was defined a unilateral boundary condition in the axial pulling out direction, local Z direction, simulating the support of the mandrel nosecap to the plate. The same boundary condition was applied to the conical part face of the split sleeve, in which the top of the nose cap avoids the pull trough of the sleeve during mandrel driving. The mandrel degrees of freedom were fixed in the x-y plane on the two top surface, and the total translation displacement of 25 mm was imposed only to the face in the nosecap side. Contacts between the three components were defined on all the facing surfaces as symmetric standard contacts with a normal penalty stiffness of 0.1 and a friction coefficient of 0.1. The mandrel was realized in a high strength steel and the split sleeve is a stainless steel tubular piece, both modeled with an elastic material definition. The implemented plate material model was the modeFRONTIER optimized Chaboche for the 7075-T73 aluminum alloy. The non-linear analysis was subdivided into 8 load steps, in order to gradually driving the mandrel in the critical contact positions: 1) first tapered part of the mandrel in contact with the sleeve, 2) mandrel maximum diameter at the entry hole side, 3), 4), 5) three equally spaced positions in the plate thickness, 6) mandrel maximum diameter at the exit hole side, 7) complete exit of the mandrel. During the last load step 8) the element components relative to the sleeve, the mandrel and the inner partition at the hole edge were removed by the ANSYS command “EKILL” to simulate the removing of the mandrel and the sleeve in the final part of the process and the reaming of the hole at the diameter 4.83 mm. The thickness of the FEM reaming had been previously calculated by the results of a simple FEM analysis in which the hole expansion and the removal of the mandrel were simulated by imposing

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Aerospace, Automotive and Marine Applications 6 96 and removing an uniform radial displacement on the nodes at the hole edge. That was to take into account in the complete analysis the radial plastic deformations of the plate that obviously change the initial dimensions of the hole. A converged solution at the highly non linear problem was obtained with a large number of sub-steps and a sparse direct solver with the Newton-Raphson option.

4. Analysis Results & Discussion

The employment of a simple axial-symmetrical analysis to investigate the process can be useful to obtain a general description of the residual stress field around a worked hole but looses important information and cannot puts in evidence some side effect of the process. The simulation of the actual mandrel drawing highlights stress and deformations during the different steps of the process. When the mandrel start to cross the sleeve the gradual opening of the split follow correctly the movement of the maximum diameter of the mandrel, as can be see in fig. 17. When the plate expansion starts it’s possible to record the evolution of the radial and tangential stress step by step along the plate thickness. In the fig. 18 and 19 are reported as example the stress histories in the material surrounding the hole edge along a radial path on the plate exit face. The three dimensional nature of the process introduce a different stress field along plate thickness. Also after the final reaming, as show in Fig. 20, the tangential component is minimum at the entry face, afterwards it increases up to the absolute maximum value at almost half of the plate thickness and slightly decreases at the exit face. The residual stress obtained can justify the different rate of propagation observed in the previous tests for defects positioned at the entry and the exit face of the treated specimen. What is more, as observed in the fracture surface of a broken specimen, insert in Fig. 20, the P shape of the crack front can follow the negative field along the thickness. The final reaming of the hole is introduced into the process to correctly size the hole and to remove possible scratches on the hole surface that could generate premature fatigue failures. In fact the introduction of the lubricated split sleeve, that mediates the contact between hole and mandrel introduce a loss of axial-symmetry. In particular, in terms of deformations, the presence of the split produces, on the hole edge, a region subjected to a no uniform radial expansion, usually termed as ‘pip’. This aspect is correctly verified in the FEM analysis as can be observed in Fig. 21. The loss of axial-symmetry obviously change also the residual stress field around the hole. The graph shows how the higher compressive stresses are obtained in correspondence of the position of the split and the lower in the diametrically opposite direction. A mean value is in the orthogonal direction. That is way usually is recommended of positioning the split in the external load direction.

5. Conclusions

In the present activity it has been set up an automatic methodology that allows, by means of modeFRONTIER optimization tool, a full management of the Ansys Chaboche material model parameters. The final aim is to develop a reliable model for materials that shown a plastic behavior under compression significantly different from the plastic behavior under tension. Classical models as isotropic or kinematic hardening could actually bring to a non conservative evaluation of the stress field. Obviously the negative effect on predictions of the structural behavior of high stressed structures during the operative life is not negligible. The Split Sleeve Cold Expansion process is now widely practiced to increase the fatigue life of critical holes in metallic aero-structures since it is capable of generate a beneficial compressive residual stress field around treated holes. However an incorrect evaluation of process effectiveness could make risky the introduction of cold working as design parameter. The developed complete FEM model highlighted useful information about the process, such as a loss of axial-symmetry due to the split sleeve and a variation of the residual stress field in the plate thickness due to mandrel drawing. Moreover the introduction of the optimized Chaboche plastic material model, tuned with focused tests on the 7075-T73 aluminum alloy, allowed a correct quantitative assessment of the compressive residual stress field. Despite more time consuming analysis, numerical results reliably matched fatigue and propagation tests. Thanks to modeFRONTIER matching flexibility, the present methodology can be proficiently applied to characterize material models in every FEM code environment. Desired extra objectives could be easily introduced to further improve methology accuracy. The effect of a possible double application of the cold expansion

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Aerospace, Automotive and Marine Applications 7 97 process will be investigated both in opposite and concord directions. An additional improvement in the plate thickness tangential residual stress distribution is expected. Next goal will be the introduction of the residual stress field in provisional software to correctly evaluate fatigue life of open hole specimens in order to consecutively extend the results in more realistic joint structures.

Acknowledgements

We would like to express our gratitude to the following people for their support and assistance in developing this research: Prof. Agostino Lanciotti, Dr. Roberto Gonella, Dr. Roberto Merlo, Dr. Enrico Troiani. We would also like to thank Dr. Armin Fritsch, involved in the initial discussions of this project. We wish to express our sincere appreciation to Prof. J. L. Chaboche for his kind help and his helpful comments.

References

[1] ModeFRONTIER® user manual 3.1.3.

[2] Ansys® Theory User Manual, version 10.0.

[3] L. Boni, A. Lanciotti, C. Polese "Residual stress field in an open hole specimens subjected to “cold working” and its effect on the fatigue resistance". Department of Aerospace Engineering at University of Pisa and Agusta Westland Research Collaboration. Documents of the Department of Aerospace Engineering at University of Pisa DDIA 2006/4, February 2006.

[4] Phillips, J. “Fatigue improvement by sleeve cold working”, SAE preprints n. 730905, 1973.

[5] Cook R. “Fatigue enhancements due to cold expansion in a range of aluminium alloy”. International Committee on Aeronautical Fatigue, 18th Symposium, Melbourne, 3-5 May 1995, pp. 971-986.

[6] Zhang. X., Wang. Z. “Fatigue life improvement in fatigue-aged fastener holes using the cold expansion technique”. Inter. Journal of Fatigue, vol. 25, n. 9-11, September/November, 2003, pp. 1249-1257.

[7] J. L. Chaboche “Unified model of cyclic viscoplasticity based on non linear kinematic hardening rule”. Handbook of Material Behavior Models, Vol. 1, Ch. 5.8, pp. 358-367, Academic Press, 2001.

[8] Ball D. L. “Elastic-plastic stress analysis of cold expanded fastener holes”. Fatigue and Fracture of Engineering Materials & Structures, vol. 18, n. 1, January 1995, pp. 47-63.

[9] Odzemir A. T., Cook R., Edwards L. “Residual stress distribution around cold worked holes”. Inter. Committee on Aeronautical Fatigue, 17th Symposium, Stockholm, 9-11 June 1993, pp. 207-237.

[10] Pavier M. J., Poussard C. G. C., Smith D. J. “Finite element simulation of the cold working process for fastener holes”. Journal of Strain Analysis for Eng. Design, vol. 32, n. 4, July 1997, pp. 287-300.

[11] Stefanescu D. “Experimental study of double cold expansion of holes”. Journal of Strain Analysis for Eng. Design, vol. 38, n. 4, July 2003, pp. 339-347.

[12] Bernard M., Bui-Quoc T., Burlat M. "Effect of re-cold working on fatigue life enhancement of a fastener hole". Fatigue Fracture Eng. Mat. Struct., vol. 18, n. 7-8, Jul-Aug 1995, pp. 765-775.

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Aerospace, Automotive and Marine Applications 8 98

Fig. 1 – Split Sleeve Cold Expansion process.

6.35 a

25.4

0

0 2 t=2.3 mm =

R b GRIP AREA 25

= 4.83 270

Fig. 2 a) Round un-notched specimen, b) open hole specimen. 9 99

600 S =465 MPa 1 500 y

400 3 300

S, MPa 2 200 2 3 100 1 0 S =215 MPa -100 y

-200 4 -300

-400 Kinematic hardening -500 Isotropic hardening -600 -0.05 -0.04 -0.03 -0.02 -0.01 0.00 0.01 0.02 0.03 0.04 0.05 0.06 ε, ΔL/L

Fig. 3 – Stress – strain curves of 7075-T73.

200

190

180

170 >

160

Smax, MPa 150 >

140 Pristine CW 130 DCWopp DCWconc Mat. 7075-T73; t=2.3 mm; R=0.1 120 104 105 106 107 Cycles, N

Fig. 4 – Fatigue test results of open hole specimens. 10 Smax=155 MPa, R=0.1 10 100 Base (No CW) CW (entry face) CW (exit face) DCWopp (entry face 1) 8 a c DCWopp (exit face 1) DCWconc (entry face) DCWconc (exit face) 6 c, mm 4

2

0 100 1000 10000 100000 1000000 cycles Fig. 5 – Crack propagation test results.

Fig. 6 – FEM model for material optimization.

11 101

Fig. 7 – Chaboche option in ANSYS material models.

Fig. 7 – Chaboche option in the Ansys FEM code

Fig. 8 – Mode Frontier work flow. 12 102

Fig. 9 – Mode Frontier DOE designs table.

Fig. 10 – Mode Frontier optimization flow. 13 103

Fig. 11 – Mode Frontier designs space.

600

400

200

0 , MPa σ -200

-400 exp. material 7075-T73 -600 Ansys Chaboche

-0.02 -0.01 0.00 0.01 0.02 0.03 0.04 0.05 ε, (ΔL/L)

Fig. 12 – Optimized FEM and experimental stress-strain curve.

14 104

Fig. 13 – ModeFrontier correlation matrix.

Fig. 14 – ModeFrontier Student Chart.

15 105

Fig. 15 – Mandrel 4-4-N and split sleeve.

a

b

Fig. 16 - Mesh a) plate-mandrel group b) plate sleeve group. 16 106

Fig. 17 – Two mandrel driving steps and the final reaming: a) entry side, b) exit side.

17 107

Fig. 18 – Tangential Stress history on the plate exit face.

18 Fig. 19 – Radial stres history on the plate exit face. 108

Fig. 20 – Tangential residual stress along plate thickness.

0 A Sleeve split B Opposite split C Split 90° -100 , MPa y

-200

A

-300

Tangential Stress S Stress Tangential C

B -400

0.00.51.01.52.02.5 x, mm Fig. 21 – Stress distribution changement due to the split sleeve. 19 109

Numerical and experimental assessment of fatigue behaviour of “cold worked” open holes

L. BONI, A. LANCIOTTI, C. POLESE

Department of Aerospace Engineering, University of Pisa, Pisa, Italy

ABSTRACT

The present paper describes the research activity carried out at the Department of Aerospace Engineering of the University of Pisa (DIA) concerning the influence of “cold working” on the fatigue resistance of open holes. The experimental activity concerned both fatigue and crack propagation tests, carried out on specimens in 7075-T73 aluminium alloy, thickness 2.3 mm, having a central hole of 4.83 mm diameter; the specimen was a simple “open hole”, i.e. no rivet or bolt were inserted. A remarkable increment of fatigue life due to the “Split Sleeve Cold ExpansionTM” process was observed. Even greater improvements were observed by applying the cold expansion twice on the same hole or by the “StressWaveTM” method. Quantitative assessments of compressive residual stress fields around cold worked holes were obtained by ABAQUS® finite element analyses. In particular, the three dimensional nature of the cold expansion process was accounted for by considering 3-D models. The cold working process was simulated both by means of a radial expansion of the hole edge and, more realistically, by modelling the mandrel drawing across the hole. The latter type of analysis highlighted an important difference between stress fields at the mandrel ‘entry’ and ‘exit’ faces, i.e. the surfaces of the specimen where the mandrel, respectively, enters and leaves the hole. The stress field numerically evaluated was confirmed experimentally by the destructive “Sachs” method. AFGROW software was consequently used to predict crack nucleation and propagation in cold-worked holes but thus far, predicted fatigue life mismatched with experimental evidence; however, numerical results were qualitatively in good agreement with fatigue tests data. Several peculiar effects of the process, experimentally observed, were clearly confirmed by FEM analyses that demonstrate to be once more an effective design tool.

1. INTRODUCTION

Cold expansion of fastener holes in metallic aircraft structure, usually known as ‘cold working’, has become a widespread practice over the last 20-30 years, mainly used in in-service aircraft to recover manifested fatigue problems or, less commonly, used in new aircraft production in order to guarantee that operational and economic lives of critical components meet or exceed required design life. The cold working process produces in the fatigue critical region a self-equilibrating residual stress field, compressive in the most critical point, which superimposes tensile stress introduced by fatigue loads. So a lower mean stress value is obtained, with evident improvement from the fatigue resistance point of view. While a number of cold expansion processes have been developed over the last years, the ‘Split Sleeve Cold ExpansionTM’ process is certainly the most successful in aeronautical industry [1]. This process consists of placing a longitudinally split sleeve within an hole and drawing an oversized tapered mandrel through this assembly, as schematically shown in fig. 1. Hole is expanded to an extent sufficient to cause permanent plastic deformation. After mandrel removing, surrounding elastic material attempts to recover its un-deformed state, producing a self-equilibrating residual stress field. Above all, a compressive circumferential residual stress is achieved in an annular region adjacent to the hole.

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During expansion phase, the sleeve avoids the direct sliding of the mandrel on the hole and, consequently, precludes hole damage that can trigger off fatigue failures. Furthermore, the sleeve limits axial displacement of the material, that, otherwise, might be subjected to a like an extrusion process. In order to reduce the mandrel pull force, the sleeve is lubricated and longitudinally split; the region at the edge of the hole subjected to a non-uniform radial expansion due to the sleeve split is usually termed as ‘pip’. This small step and possible scratches require an hole final reaming. Appropriate and calibrated tools must be obviously used to perform all the described operations. According to the previous considerations, the split-sleeve cold expansion of fastener holes generates a complex, three dimensional residual stress field. In particular, such a residual stress field varies through the thickness of the component and is asymmetric around the hole. The residual stress field is not axial-symmetric because of the presence of the split in the sleeve: the maximum circumferential stress is reached in correspondence of the split, and the minimum is on the diametrically opposite location. As a consequence, fatigue resistance depends on the sleeve split position respect to the load direction, [2]. Usually, the best orientation of the split is coincident with the load direction, so having the same conditions in both sides of the hole. A number of studies on the cold working technique have been performed by simulating the process as an uniform radial expansion of the hole, i.e. by accepting the hypothesis of axial- symmetry of the problem. These models obviously cannot account for the split of the sleeve and the material flow in axial direction that is still appreciable, even in presence of the sleeve. Experimentally, strain gauge measurements or, better, X-ray diffraction measurements have shown remarkable differences between the residual stress field at the mandrel entry and exit faces. In particular, the residual stress is higher at the exit face than at the entry face, that is consistent with the faster crack growth at the inlet face with respect to the outlet face, observed in several experimental investigations, [3, 4, 5]. Finite Element analyses, performed by simulating the mandrel drawing across the hole confirm the differences between the stress states at the mandrel entry and exit faces, [6]. This suggests that double cold working in reverse directions may be a solution for structures for which there is access from both sides. In particular, as reported in [7], the application of such a technique produces further beneficial effects on the fatigue life with respect to the standard single cold working; the author justifies this result by considering that the double cold working reduces the difference between residual stresses on the inlet and outlet faces introduced by the material flow in axial direction: the residual interference measured at the entry face is about 60-80% or 95-96% of the one measured at the exit face for the case of simple and double cold working respectively. In [8] it is shown that also a repeated application of the cold expansion in the same direction improves the fatigue life, in contrast, at least in part, with the explanation reported in the previous work relevant to the effect of the material flow in axial direction. As a matter of fact, the double cold working performed in the same direction is very interesting since it can be applied also to the structures for which the access is allowed only from one side. Another difficulty intrinsically connected with the modelling of the cold working process is represented by the material behaviour, since it can reach, after the mandrel extraction, as a function of the adopted level of interference, elastic-plastic deformations during the unloading phase, by causing a second plasticization of the material around the hole. This phenomenon can be anticipated by the Bauschinger’s effect. All the described problems must be faced by means of expensive experimental tests, so that the use of cold working is basically limited to the solution of fatigue problems encountered in service. A deeper knowledge of such problems could allow to further spread this technique. Split sleeve cold working is the most known system to produce compressive residual stress field, but it is not the only one. StressWaveTM is a relatively new method, completely different from the previous one, which introduces compressive residual stresses prior to machining the hole, without any pre- or extra post-processing operations. The method, developed by StressWave Inc. [9], uses hardened indenters to produce plastic deformation in an area smaller than the diameter of the hole that will be subsequently drilled, fig. 2. Witnesses of this process phase are small dimples on both side of the work piece. In this case uniform residual stresses

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are introduced in radial direction. Hole drilling is the second step of the process. The entire dimple and the surface upset are also removed. The process has been applied in a variety of alloys (aluminium, steel, titanium and cast iron) and section thickness (0.8 to 25 mm). Due to the great industry interest on these processes, an experimental and numerical research started at the Department of Aerospace Engineering of Pisa of the University of Pisa, aimed at the evaluation of effectiveness of cold working techniques for their application in aeronautical structures, [10]. Open hole specimens were fatigue tested under constant amplitude loading to investigate different aspects of cold working processes.

2. MATERIALS, GEOMETRY AND PREPARATION OF SPECIMENS

A typical aeronautical aluminium alloy 7075-T73 was examined, with longitudinal grain direction in all the specimens. Static tests were performed on specimens with a circular cross- section of 6.35 mm diameter and a central cylindrical part of 25.4 mm length, specimen A in fig. 3a, machined from a plate of 15 mm thickness. Such specimens can be stressed in compression beyond the yield point without instability phenomena occur. Additional static and fatigue tests were performed on un-notched specimens, thickness 2.3 mm, width 10 mm, specimen B in fig. 3b, to dispose of all the information necessary for fatigue life prediction of cold worked specimens. Test on cold worked open hole were carried out on dog bone specimens, thickness 2.3 mm, with a final reamed diameter of 4.83 mm, specimen C in fig.3c. Split sleeve cold expansion preparation was realized at DIA using an FTI 4-4-N mandrel, with a starting hole of 4.3 mm. In all the specimens cold working was carried out by positioning the split sleeve in the load direction. StressWave preparation was instead realized at StressWave Inc. in Seattle. So five different groups were tested: OH, specimens having only an untreated open hole, CW, with holes cold worked by FTI instrumentation and DCW specimens with a double cold working in opposite and concordance directions and CW-SW specimens cold worked by the StressWave method. The identification code of the open hole specimens is given in tab. I. In propagation tests a starting artificial flaw was introduced by blade cut; the flaw was oriented perpendicular to the load direction and was present only on one side.

Spec. Ident. Cold working Notes OH no Open hole CW yes Open hole. Split sleeve cold working CW-SW yes Open hole. StressWave DCW-opp yes (x2) Open hole. Double cold working in opposite directions DCW-con yes (x2) Open hole. Double cold working in concord directions

Tab. I - Identification of specimens. Material 7075-T73, thickness 2.3 mm.

3. STATIC AND FATIGUE TEST RESULTS OF UN-NOTCHED AND OPEN HOLE SPECIMENS

3.1 Stress-strain curve of 7075-T73

In order to describe the total stress-strain curve of 7075-T73, four A specimens were statically tested. Fig. 4 shows the results of the material stressed in tension until failure (labelled in Fig. 4 as 1); in the same figure are shown also the curves relevant to specimens stressed in tension until a total strain of about 3% and 4% and successively stressed in compression until failure (respectively labelled as 2 and 3) and the curve relevant to a specimen stressed in compression until failure (labelled as 4). In the same graph, the curves relevant to the most common models

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for the material plasticity are also reported, kinematic and isotropic hardening models. It is evident that the examined material behaviour agrees fairly well with the kinematic hardening model, although remarkable differences are present between the behaviour in tension and in compression. The yielding in tension occurs quite abruptly, while the compression branch exhibits a significantly more gradual behaviour, more evident for the case of higher pre- tensioning (curve 3). The remarkable decrease of the yielding stress in compression following tension yielding attests a strong ‘Baushinger’s effect’ of the material.

3.2 Fatigue tests on 7075-T73 un-notched specimens

In order to better understand the fatigue behaviour of the material under examination, fatigue tests were carried out on un-notched specimens B. All fatigue tests were carried out at a stress ratio, R=Smin/Smax, equal to 0.1. Some specimens were pre-stressed before fatigue tests, to simulate the strain history at the edge of a cold worked hole. Pre-stressing was carried out under strain control up to a 4% total strain. Pre-stressing shows to have a slight detrimental effect from the fatigue point of view.

3.3 Fatigue and crack propagation tests on open hole specimens

Fatigue tests were carried out on open hole specimens C. The results obtained, shown in fig. 5, are relevant to specimens with pristine and cold expanded holes. At higher number of cycles, the results show an increase of about 11% of the allowable fatigue stress for the case of conventional split sleeve cold working, while such an increment reaches about 18% and 25% for the case of double cold working performed in concord and opposite directions respectively. Obviously, due to the characteristic shape of fatigue curves, the improvements introduced by cold working are more evident in terms of fatigue life. StressWave effect is comparable with double split sleeve cold working. The crack propagation problem was investigated separately, by performing comparative tests at two stress levels (maximum stress of 155 MPa and 175 MPa R=0.1) on specimens having an initial corner defect of about 0.1 mm length in the two directions (width and thickness, as shown in the inset of fig. 6). The artificial defect, produced by means of a cutter blade, resulted to be very effective: in all tests performed at a maximum stress of 155 MPa, a fatigue crack nucleated after 5000÷10000 cycles. The results obtained relevant to four conditions of pristine specimens and specimens subjected to simple split sleeve cold working, double cold working in opposite and concord directions are compared in fig. 6. An initial surface defect size equal to 0.3 mm was adopted; actually even smaller cracks were identified, but the value c0=0.3 mm was present in each test and it is large enough to be not influenced by the initial artificial flaw. The two cases of initial defect lying on the entry and the exit face were considered. In the specimens subjected to double cold working in opposite directions, the entry face is the one relevant to the first mandrel drawing. It is evident the strong retardation effect of the cold working on the crack propagation, fig. 6. In particular, for a no-cold worked hole, an initial defect of 0.3 mm becomes critical after about 10000 cycles (Smax=155 MPa, R=0.1). In the same conditions, the life of cold worked specimens is 10 and 30 times greater than the one of pristine specimens, for the cases of initial defect on the entry and the exit face respectively. This behaviour confirms the different effectiveness of the cold working at the entry and exit sides of the mandrel. In the case of double cold working performed in opposite directions, the life increases of about 60 times for the entry face and of more than 100 times for the exit face. A second application of cold working in reverse direction is thus able to improve the fatigue resistance on both sides. A crack on the exit side of holes double cold worked in concord directions has a fatigue life slightly higher than on simple cold worked holes. On the contrary, a crack on the entry side of the mandrel, once it reached, with a decreasing growth rate, a size of 0.6 mm in about 700000 cycles not further

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propagated, until test end at more than 2.5ּ106 cycles. Indeed experimental results confirm the positive effect of double cold working performed in concord direction; it is however conceptually difficult to find a justification for such a behaviour. Crack propagation tests were carried out also on StressWave cold worked specimens. As explained previously, crack propagation on pristine and split sleeve cold worked specimens were carried out at Smax=155 MPa, R=0.1. When this fatigue stress was applied to StressWave cold worked holes, cracks nucleated but do not propagated. Fatigue cracks stopped when crack length was about 1 mm. The same results were obtained when higher fatigue stresses, 160 and 165 MPa, were applied. Crack propagated up to the final failure of StressWave cold worked specimen when fatigue stress was increased to Smax=175 MPa. Crack propagation life of StressWave cold worked specimen is very high when compared with pristine specimen, fig. 7, even if a second fatigue crack nucleated during fatigue test. An approximate comparison between crack propagation test results of Split Sleeve and StressWave cold worked specimens is presented in fig. 8. The results can be compared when transformed in the form: crack propagation rate as a function of stress intensity factor range; in the present case the analysis is approximate because the actual shape of the cracks is un- known, such as consequently the Stress Intensity Factor. Cracks nucleated as corner cracks and propagated with an oblique front; only crack surface dimension was monitored during the tests. To compare the results, it was assumed that cracks were through cracks; consequently stress intensity factor is over-estimated. Anyway, it is reasoning to assume that the results relevant to the different cold working systems are affected by the same errors. With the approximations described, the results shown in fig. 8 indicate a strong reduction of crack propagation rate in cold worked holes respect to pristine holes at low ΔK values. At higher ΔK values crack propagation rate in Split Sleeve cold worked holes increases up to values measured in pristine holes, while the beneficial effect of StressWave cold working is present in a larger area.

4. NUMERICAL/EXPERIMENTAL EVALUATION OF THE RESIDUAL STRESS FIELD OF THE MATERIAL SURROUNDING A COLD WORKED HOLE

To better understand the experimental results, a number of numerical simulations of the cold working process were performed, with the aim of investigating the contribution of the parameters involved in the process itself. In particular, several analyses were carried out, characterised by different degrees of approximation in the process simulation, all performed using the Finite Element code ABAQUS 6.4, [11], which proved to be a very versatile software. The three dimensional nature of the cold expansion process was accounted for by considering 3-D models. In parallel with the numerical approach, an experimental evaluation of the residual stress field around a cold worked hole was performed by using the Sachs method [12]. Finally, the numerical/experimental results obtained were used to predict the fatigue life of components having holes subjected to cold working; as it will be explained in the following, at the moment, this problem seems to be far from the solution.

4.1 Description of Finite Element models: geometry and material

For each FEM analysis, the compressive residual stress field was estimated in the material surrounding the hole of a circular plate with a 50 mm diameter and a 2.3 mm thickness. The diameter of the plate was selected to be large enough respect to the hole diameter. It was demonstrated that the residual stress field is not significantly different for smaller plates, having the diameter equal to the width of the fatigue specimens, i.e. 25 mm. The circular plate was made of 7075-T73 aluminium alloy, while both mandrel and split sleeve built in steel. The material data relevant to the 7075-T73 aluminium alloy were experimentally determined and related to the stress-strain curve shown in fig. 4 and previously described. Actually, the same behaviour under loads in tension and compression was

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supposed, since different properties of the material in tension and compression cannot be effectively assigned in the FEM package used, [11]. According to the better agreement with the experimental results, the kinematic plastic hardening material model was selected.

4.2 Split sleeve cold expansion FEM models

The FTI specifications were applied to an hole with an initial diameter of 4.3 mm, relevant to an FTI 4-4-N mandrel and to a split sleeve of thickness 0.152 mm. The reaming of the hole that brings the hole diameter to the final size of 4.83 mm was also considered.

4.2.1 Types of analysis

The FTI specifications for the cold expansion process can be divided into three main steps, which lead to develop the residual stress FEM model: 1) Hole expansion by pulling a tapered mandrel through the hole. 2) Hole unloading once the mandrel is removed. 3) Finish ream that removes a thin layer of material around the hole. In the actual process, the presence of a lubricated split sleeve must be accounted for, interposed between the hole edge and the mandrel surface. The difference among the considered analyses is relevant to the first two phases, while the phase 3 is in any case simulated by the function ‘element removal’, available in ABAQUS 6.4. As far as the phases 1 and 2 are concerned, the different kinds of analysis that were performed can be summarised as follows:

ANALYSIS 1

- The hole expansion (phase 1) is simulated by imposing an uniform radial displacement on the nodes at the hole edge. This is the simplest technique to simulated cold working. The mandrel is supposed rigid and consequently is not modelled. The radial displacement imposed is equal to half difference between the diameters of mandrel plus the sleeve thicknesses and the hole. - The removal of the mandrel (phase 2) and the corresponding unloading process are simulated by eliminating the boundary condition at the hole edge. - The presence of the split sleeve is not accounted for.

ANALYSIS 2

- The mandrel, including sleeve thicknesses, is directly modelled and its drawing across the hole, from the entry to the exit face of the plate, is realistically simulated, by imposing to it a translation displacement. The analysis is significantly more complex than the previous one. - The surface-to-surface contact interaction is applied between the mandrel and the hole surfaces, by supposing a pure normal behaviour for the contact.

ANALYSIS 3

- The split sleeve is modelled. - The hole expansion (phase 1) is obtained by imposing an uniform radial displacement to the nodes of the internal surface of the sleeve. - The hole unloading (phase 2) is simulated by removing the radial displacement on the sleeve internal surface.

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- The contact between the hole edge and the external surface of the sleeve is modelled by the “surface-to-surface” contact interaction. Both the normal and the tangential contact behaviour are defined, with a supposed value of the friction coefficient equal to 0.1.

Furthermore, starting from the configuration relevant to ANALYSIS 2, where the mandrel is physically modelled, others two analysis (in the following indicated as ANALYSIS 4 and ANALYSIS 5) were performed to investigate the double cold working. A second mandrel was added, crossing the hole in the opposite direction with respect to the first one in ANALYSIS 4 and in the concord direction in ANALYSIS 5. Tab. II gives a synthesis of the analyses performed.

ANALYSIS 1 Hole expansion ANALYSIS 2 Mandrel drawing across the hole ANALYSIS 3 Sleeve included in the model ANALYSIS 4 Double cold working in opposite directions ANALYSIS 5 Double cold working in concord directions

Tab. II - Analyses performed

4.2.2 Mesh

The main characteristics of the meshing technique, common to all the above described models, are: - The 8-node 3-D solid element was adopted. - For the plate, 76, 26 and 7 elements were used, respectively, in circumferential, radial and thickness directions; in the radial direction a mesh seed was applied in order to obtain a higher density of elements close to the hole edge. In fig. 9 is shown the mesh relevant to the groups plate-mandrel and plate-sleeve.

4.2.3 Boundary conditions

According to the split sleeve cold expansion system proposed by the FTI, during the mandrel pulling, a nosecap with circular cross-section, connected to the puller unit, is held firmly against the work piece. In all the models the displacement in the mandrel pulling direction was restrained on the nodes lying on a circumference on the exit face, coaxial to the hole. Actually, the constrained nodes are located not too far from the hole edge, to realistically simulate the constraint introduced by the nosecap, but far enough from the hole edge to avoid negative interaction effects. Particularly, the constrained nodes are distant 3 mm from the hole edge, lying in a region where the residual stress field is quite stabilized, by reaching asymptotic values. Displacements of the plate in its plane were avoided by imposing translation constraints on two opposite generatrices. Analogous translation constraints were applied on the centre of both the mandrel base faces, leaving free the displacement in the drawing direction. As far ANALYSIS 3 is concerned, the translation degrees of freedom in circumferential and axial directions were constrained along the generatrix of the sleeve external surface opposite to the split.

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4.2.4 Results of the analyses

ANALYSIS 1

In fig. 10 are shown the residual stresses both in tangential and radial direction, relevant to a generic radial path from the hole edge towards the plate periphery. The model analysed is axial-symmetrical and, due to the type of cold expansion simulation, it does not exhibit any difference between entry and exit face. The residual stresses, plotted as a function of the distance from the hole edge are reported before and after the reaming of the hole. It is possible to see that the compressive stress field in the material surrounding the hole is mainly due to the tangential component. Furthermore, it is interesting to observe (see the zoom in fig. 10) that the finish reaming removes the small area of reversed yielding that characterises the region close to the hole edge. The beneficial effect of hole reaming, which causes larger compressive residual stresses at the hole edge, is attested also in [5]. The radial stress reaches its maximum value, about 50% of the maximum tangential stress, at about 2 mm from the hole boundary. Since the variation of the radial stress in the different performed analyses is quite limited and less interesting with respect to the longitudinal residual stress, this information will not be given in the following discussion.

ANALYSIS 2

The simulation of the actual mandrel drawing introduces a variation of the residual stress field in the plate thickness, by preserving the axial-symmetry of the model. By considering fig. 11, where the residual stresses are represented along the hole axis, following the direction from the entry to the exit face, the three dimensional nature of the cold working process is highlighted: the tangential component is minimum at the entry face, afterwards it increases up to an almost constant value from one third of the plate thickness until the exit face, reaching the absolute maximum value at about the middle distance. Comparing present results to those relevant to ANALYSIS 1, it can be observed that, in the previous model, the tangential residual stress is overestimated at the entry face while at the exit face it is slightly underestimated.

ANALYSIS 3

The aim of this analysis is to investigate the loss of axial-symmetry introduced by the action of the lubricated split sleeve, that mediates the contact between hole and mandrel. In particular, in terms of deformation, the presence of the split produces, on the hole edge, a region subjected to a non-uniform radial expansion, usually termed as ‘pip’, visible in fig. 12. In terms of residual stress field, in the graph of fig. 13, a comparison is shown between the results of the model without sleeve (ANALYSIS 1) and those relevant to ANALYSIS 3, reported for the locations A, B and C specified in the figure. The split sleeve introduces a slight asymmetry, mainly evident at the hole edge. Indeed, according to the represented results, the most compressed region is the one lying on the split side only for the material at a certain distance from the hole edge: the region close to the hole edge undergoes a significant reversed yielding, that alters the described behaviour. Actually, the entity and the region of extension of the reversed yielding phenomenon might be overestimated by the kinematic model, adopted to describe the plastic hardening of the material. As a matter of fact, according to measurements performed in [13], the maximum compressive residual stress is reached in correspondence of the split of the sleeve.

ANALYSIS 4 & 5

The effect of a double application of the cold expansion process was investigated by means of this model, accounting also for the case of reverse direction of the mandrel drawing. By observing the graph shown in fig. 10, where the results relevant to simple cold working (ANALYSIS 2) and double cold working (ANALYSIS 4) are compared, one can see that the

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compressive residual field at the entry face clearly improves, while a worsening results at the exit face. The 3-D behaviour of the residual stresses shows how the improvement, in terms of tangential component, characterises about the four fifth of the plate thickness, while the region close to the exit face gradually worsens. The results of double cold working in concord directions (ANALYSIS 5), given in fig. 10, are practically coincident with the results of standard cold working; a justification of the beneficial effect of double cold working in concord directions, experimentally observed, was not found numerically.

4.3 StressWave cold working Fem model

StressWave Inc. developed a Stepped-type indenter to be applied to the same final hole diameter than the split sleeve cold expanded holes, d=4.83 mm. The geometry of the S-type indenters and the compressive target loads were selected in order to obtain a competitive compressive residual stress field for this application. Indenter shape details are proprietary information of StressWave Inc..

4.3.1 FEM analysis

An axial-symmetrical model of the circular plate was developed. The stepped indenter developed by StressWave Inc. was modelled as elastic steel parts with an highly simplified geometry. The symmetry respect to half thickness plate allowed to reduce the complexity of the problem to half model. As before the kinematic plastic hardening material model was selected for the 7075-T73 aluminium alloy plate. The contact between the indenter and the plate was modelled by the “surface-to-surface” contact interaction. Both the normal and the tangential contact behaviour were defined, with a supposed value of the friction coefficient equal to 0.1. The StressWave cold working process was simulated imposing an axial displacement to the indenter and after removing it, fig. 14. The hole in the final hole configuration was obtained by the function ‘element remove’, available in ABAQUS 6.4.

4.3.2 Mesh

The main characteristics of the meshing technique are: - The axial-symmetric 3-D CAX8R solid element was adopted. - A uniform, quite high density mesh seed was applied thanks to model simplicity in order to obtain a higher definition of the problem, fig. 14.

4.3.3 Boundary conditions

An axial symmetry and a middle-plate thickness symmetry were introduced. The plate was restrained on the nodes at the external face, far enough from the hole edge to avoid negative interaction effects.

4.3.4 Results of FEM analysis

The first evaluation of FEM analyses can be obtained comparing a section of a StressWave cold worked plate with the plastic deformation present after indenter removing, fig. 15. A good agreement in the residual dimple depth and geometry can be observed, despite high simplification of the problem. Compared with the split sleeve cold working, fig. 12, a slightly

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higher compressive residual stress is obtained and a significant reverse yielding can be observed after the drilling of the hole, fig. 16. The region around the hole with beneficial compressive residual stress is approximately of the same dimensions than in the split sleeve cold working, but the stress field is constant along plate thickness.

5. EXPERIMENTAL EVALUATION OF THE RESIDUAL STRESS FIELD

An experimental evaluation of the residual stress field around a cold worked hole was carried out by the Sachs method, [12], a widely used instrument that despite its simplicity well compare with more sophisticated methods, such as synchrotron X-ray diffraction and neutron diffraction techniques, [14] . According to this destructive technique, material is removed layer by layer from the bore of an axial-symmetrical component and the residual stresses present prior to boring are obtained by the elastic relationship between the exterior surface deformations and the stress existing in the removed material. Experimental measurements were carried out in two circular plates, diameter 25 mm, in 7075-T73 aluminium alloy, thickness 2.3 mm, with a central cold worked hole, diameter 4.83 mm after reaming; the split of the sleeve was located at 90° with respect to the position of the strain gauges. Due to the small thickness of the plates, a plane stress state can be hypothesized, so that the equations of the Sachs method can be simplified and only one strain gauge in circumferential direction is necessary to obtain both radial and hoop stress, according to the following relationships [15]:

d ⎡ ⎛ r 2 − r 2 ⎞⎤ σ = E ε ⎜ e ⎟ h ⎢ h ⎜ 2 ⎟⎥ dr ⎣ ⎝ 2r ⎠⎦

⎡ ⎛ r 2 − r 2 ⎞⎤ σ = E ε ⎜ e ⎟ r ⎢ h ⎜ 2 ⎟⎥ ⎣ ⎝ 2r ⎠⎦

where σ h and σ r are the hoop and radial stress, ε h is the hoop strain, E is the Young’s

modulus, re is the external radius and r the internal radius, which is progressively increased during the measure. The results were obtained according to the procedure suggested in [16] for the evaluation of the hoop stress derivative. The experimental test and the numerical analysis give rise comparable results, fig. 17. However it had to be highlight that the Sachs method applies elastic relationships, so it overestimates residual stress in the elastic-plastic field. Besides numerical results are affected by the hypothesis of equivalent behaviour of the material subjected to compressive or tensile stresses. In [5] the residual stress field measured after fatigue loading show a decrease (relaxation) proportional to the number of applied cycles. Consequently five specimens B containing a cold worked hole were fatigue tested, Smax=160 MPa, R=0.1, until respectively the 20, 40, 60, 80 and 100% of mean fatigue life (equal to 1260000 cycles from fig. 5). After the five specimens were milled to obtain circular plates, external diameter 24 mm, coaxial with the cold worked hole. Measurement by the Sachs method did not show significant modification of the residual stress field as a function of the number of applied fatigue loads. Residual stresses were measured also in two StressWave cold worked specimens, fig. 17. Also in this case, the Sachs method and the simulation by ABAQUS of the same method give rise to comparable results.

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6. PREDICTION OF FATIGUE LIFE OF COLD WORKED HOLES

Even if the natural application of the measurement and/or the evaluation of the residual stress field around a cold worked hole would have to be finalized to the fatigue life evaluation, in actual fact, only a few papers deal with this problem. In the present case the basic idea was the introduction of the residual stresses in a pre-existing fracture mechanics computer program; AFGROW was selected for this purposes, [17]. AFGROW can account for the existence of residual stresses by calculating additive residual stress intensity factors Kres, which are added to the stress intensities (both maximum and minimum) caused by the applied loads. This will not change ΔK, but will change the stress ratio, R. The program is not able to evaluate the redistribution of the residual stress field due to crack propagation; this point, even if important, is not fundamental, because a large part of fatigue life is consumed in presence of small cracks, which would not cause significant residual stresses redistribution. Besides, redistribution is less important in the case of compressive residual stresses, because, contrary to the case of tensile residual stresses, compressive residual stresses can be present also in a cracked body. In fig. 18 the experimental crack propagation data are compared with those predicted by AFGROW, relevant to a corner crack with initial size 0.3 mm in both directions. The experimental curves are relevant to a no cold worked hole and to cold worked holes having start defect at the entry or at the exit faces. All the represented data refers to a maximum stress Smax=155 MPa and to a stress ratio R=0.1. The predicted curve with no cold working shows a good agreement with the experimental data, while, a similar agreement can be obtained for the cold working case only by considering a small amount, about 30%, of the estimated residual stress field. AFGROW has a module, based on the "local strain approach", to calculate crack initiation life, so that the program can be used also for the evaluation of the total (initiation + propagation) fatigue life. In this case the predicted fatigue resistance of specimens with cold worked holes was enormously higher than the experimental evidence. Also In this case agreement was found by considering only 20% of the residual stresses (curve CW; AFGROW, RS 20%).

7. CONCLUSIONS

A research activity was concluded at the Department of Aerospace Engineering of Pisa, relevant to the fatigue resistance of structural components with holes expanded by means of the "Split Sleeve" and "StressWave" cold working techniques. The following conclusions were obtained: Fatigue tests were carried out on open hole specimens, being this configuration simpler than riveted joints. Residual stresses were evaluated and experimentally measured in cold worked holes. Only by considering a small amount, 20 - 30%, of the estimated residual stress field predicted fatigue and crack propagation results agree with the experimental ones. Even if a justification for this result was not identified, different tests were carried out at least to exclude some possible causes, such as residual stress relaxation due to fatigue loading.

REFERENCES

[1] Phillips, J. “Fatigue improvement by sleeve cold working”, SAE preprints n. 730905, 1973.

[2] Cook R. “Fatigue enhancements due to cold expansion in a range of aluminium alloy”. International Committee on Aeronautical Fatigue, 18th Symposium, Melbourne, 3-5 May 1995, pp. 971-986.

17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 11 120

[3] Zhang. X., Wang. Z. “Fatigue life improvement in fatigue-aged fastener holes using the cold expansion technique”. Inter. Journal of Fatigue, vol. 25, n. 9-11, September/November, 2003, pp. 1249-1257.

[4] Pell R. A., Beaver P. W., Mann J. Y., Sparrow J. G. “Fatigue of thick-section cold-expanded holes with and without cracks”. Fatigue Fracture Eng. Mat. Struct., vol. 2, n. 6, 1989, pp. 558- 567.

[5] Odzemir A. T., Edwards L. “Relaxation of residual stresses at cold-worked fastener holes due to fatigue loading”. Fatigue and Fracture of Eng. Materials & Struct., vol. 20, n. 10, 1997, pp. 1443-1451.

[6] Pavier M. J., Poussard C. G. C., Smith D. J. “Finite element simulation of the cold working process for fastener holes”. Journal of Strain Analysis for Eng. Design, vol. 32, n. 4, July 1997, pp. 287-300. [7] Stefanescu D. “Experimental study of double cold expansion of holes”. Journal of Strain Analysis for Eng. Design, vol. 38, n. 4, July 2003, pp. 339-347.

[8] Bernard M., Bui-Quoc T., Burlat M. "Effect of re-cold working on fatigue life enhancement of a fastener hole". Fatigue Fracture Eng. Mat. Struct., vol. 18, n. 7-8, Jul-Aug 1995, pp. 765-775.

[9] Eric T. Easterbrook, Brian D. Flinn, Chris Meyer and Nil Juhlin, The StressWave Fatigue Life Enhancement Process, Proceedings of the 2001 Aerospace Congress, 2001-01-2578

[10] L. Boni, A. Lanciotti, C. Polese "Residual stress field in an open hole specimens subjected to “cold working” and its effect on the fatigue resistance". Department of Aerospace Engineering at University of Pisa and Agusta Westland Research Collaboration. Documents of the Department of Aerospace Engineering at University of Pisa DDIA 2006/4, February 2006.

[11] ABAQUS Version 6.4, Copyright © Hibbit, Karlson & Sorensen, Italy.

[12] Sachs G., Van Horn K. R. “Practical Metallurgy: Applied Physical Metallurgy and the Industrial Processing of Ferrous and Nonferrous Metals and Alloys”. American Society for Metals, Cleveland OH, 1940.

[13] Boni L., Lanciotti A., “Tensioni residue in un foro trattato con cold-working e loro influenza sulla resistenza a fatica”, Proceedings of the XVIII AIDAA National Congress, Volterra, 19- 22 September 2005.

[14] Stefanescu D., Steuwer A., Owen R.A., Nadri B., Edwards L., Fitzpatrick M.E., Withers P.J. "Elastic strains around cracked cold-expanded fastener holes measured using the synchrotron X-ray diffraction technique". Journal of Strain Analysis for Engineering Design, vol. 39, n. 5, September, 2004, pp. 459-469.

[15] Weiss V. “Residual stress in cylinder”. Soc. For Exp. Stress Analysis, Vol.15, n.2, 1957, pp. 53-56.

[16] De Swardt R. R. “Finite element simulation of the Sachs boring method of measuring residual stresses in thick-walled cylinders”. Transaction of the ASME, Vol. 125, 2003, pp. 274-276.

[17] Harter J. A. “AFGROW user guide and technical manual”. Airforce Research Laboratory, v4- 0007-12-11.

17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 12 121

Fig. 1 – Split sleeve cold working

Fig. 2 – StressWave cold working.

17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 13 6.35 122 a

25.4

t=2.3 mm

b GRIP AREA 10

70 270

0

0 2 t=2.3 mm =

R c GRIP AREA 25

= 4.83 270

Fig. 3 a) Round un-notched specimen, b) flat un-notched specimen, c) open hole specimen.

600

S =465 MPa 1 500 y

400 3 300

S (MPa) 2 200 2 3 100 1 0

-100 S =215 MPa y -200 4 -300 Kinematic hardening

-400 Isotropic hardening -500

-600 -0.05 -0.04 -0.03 -0.02 -0.01 0.00 0.01 0.02 0.03 0.04 0.05 0.06 ε (ΔL/L)

Fig. 4 – Stress – strain curves of 7075-T73.

17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 14 200 123 195 190 185 180 175 170 > 165 > 160 > 155 150 > 145 Smax (MPa) 140 Pristine 135 CW 130 DCWopp DCWconc 125 Stress Wave Mat. 7075-T73; t=2.3 mm; R=0.1 120 104 105 106 107 Cycles Fig. 5 – Fatigue test results of open hole specimens.

10 Smax=155 MPa, R=0.1

8 Base (No CW)

CW (entry face) a c CW (exit face) 6 DCWopp (entry face 1) DCWopp (exit face 1) DCWconc (entry face) DCWconc (exit face)

c (mm) 4

2

0 1000 10000 100000 1000000 Cycles Fig. 6 – Crack propagation test results.

17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 15 124 10 Smax=175 MPa, R=0.1

StressWave (Experimental) 8 Plain Hole (AFGROW)

6 c1

c (mm) 4 c2 c1

2 c2

0 50000 100000 150000 200000 250000 Cycles Fig. 7 – Crack propagation in a pristine specimen (simulated by AFGROW) and in a cold worked (StessWave) specimen.

0.1 Pristine hole Split Sleeve Cold Working StressWave Cold Working

0.01

1E-3 da/dN (mm/cycle) da/dN 1E-4

1E-5

a 1E-6 10 20 30 ΔΚ (MPa, m) Fig. 8 – Crack propagation rate in Split Sleeve and Stress Wave cold worked specimens.

17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 16 125

a b

Fig. 9 - Mesh a) plate-mandrel group b) plate sleeve group.

Tangential stress Radial stress 100 Tang. stress (after reaming)

0

-100 -260

-280

-200 -300 -320

S (MPa) -340 -300 -360 S (MPa)

-380 New hole edge after reaming -400 -400

-420

-500 -440 0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 0246810 x (mm) x (mm)

Fig. 10 - Residual stresses before and after the hole reaming for ANALYSIS 1.

17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 17 126 Direction entry face - exit face -300

CW -350 DCW conc. DCW opp. SW

-400 Tangential MPa Stress,

-450 0.0 0.2 0.4 0.6 0.8 1.0 no dimensional thickness, mm/mm

Fig. 11 - Residual stresses along the thickness of the plate in standard and double Split Sleeve (opposite and concord directions) and StressWave cold working

Fig. 12 – Stress distribution change due to the split sleeve.

17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 18 127

-100 C

B

-200 A

S (MPa) -300

Radial expansion (ANALYSIS 1) SPLIT SLEEVE (C) -400 SPLIT SLEEVE (A) SPLIT SLEEVE (B)

0.0 0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6 1.8 2.0 x (mm)

Fig. 13 – Effect of the split sleeve on the residual stress distribution.

Fig. 16 – StressWave cold working process Fem model.

17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 19 2.5 128 2.0 dimple

1.5

1.0

U (mm) 0.5

0.0

-0.5 0.0 0.5 1.0 1.5 2.0 2.5 3.0 r (mm) 2.3

4.83

Fig. 15. - Section of a StressWave indentation and FEM dimple result.

before drilling 200 after reaming

100

0

-100

-200 s (MPa)

-300

-400

-500

0246810 r (mm) Fig. 16 - Residual stresses before and after hole reaming in StressWave FEM model.

17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 20 129 100

0

-100 S.G.

-200

-300 S.G. S (MPa)

-400 Sachs CW -500 FEM CW Sachs SW -600 FEM SW

0123 X (mm) Fig. 17 – Comparison between predicted and measured tangential stresses.

Smax=155 MPa, R=0.1 10 Exp. Plain hole (No CW) CW Hole (Entry face) CW Hole (Exit face) 8 Plain Hole AFGROW CW Hole AFGROW (RS=30%)

6

c (mm) 4

2

0 1000 10000 100000 Cicli Fig. 18 – Crack propagation in plain and cold worked holes; comparison with the theoretical results. 17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy 21 130 Caratterizzazione "Damage Tolerance" dell'attacco ala-fusoliera del velivolo M346

Agostino Lanciotti(1), Francesca Nigro (2), Claudia Polese(1) E-Mail: [email protected]; [email protected]; [email protected]

(1) Dipartimento di Ingegneria Aerospaziale, Università di Pisa, Via G. Caruso, 56122 Pisa (2) AerMacchi S.p.A., Via Foresio n. 1, 21040 Venegono Superiore

Sommario

Il Dipartimento di Ingegneria Aerospaziale dell'Università di Pisa sta svolgendo alcune attività a supporto della certificazione "Damage Tolerance" del nuovo addestratore, M346, della AerMacchi S.p.A. (Venegono Superiore). Fra i diversi elementi strutturali critici dal punto di vista della resistenza a fatica, particolarmente interessante è l'attacco ala-fusoliera, per il quale è stato sviluppato un programma di prove sperimentali, i cui risultati saranno di ausilio anche per la progettazione di altre parti del velivolo. Il lavoro svolto ha riguardato la caratterizzazione delle proprietà di Damage Tolerance del materiale utilizzato per la costruzione delle ordinate di forza e dei longheroni alari, lega leggera 7050-T7451; le prove sono state effettuate sia con carico sinusoidale, sia con una sequenza di carico rappresentativa dell'impiego del velivolo ed hanno riguardato due diversi tipi di elementi, un provino Compact Tension standard e un dettaglio strutturale, collegamento "Lug-Fork", il cui disegno è stato ricavato dal disegno della zona più critica nel collegamento ala-fusoliera. Nel secondo tipo di provino sono state installate anche delle bussole ad alta interferenza, ForceMateÒ, che si sono dimostrate molto efficaci dal punto di vista della resistenza a fatica. Su alcuni provini sono stati introdotti dei difetti artificiali, sia prima che dopo l'inserimento delle bussole. Il lavoro, ancora di fase svolgimento, ha riguardato anche alcuni aspetti numerici, mirati alla valutazione dello stato di tensione prodotto dall'installazione delle bussole.

INTRODUZIONE

Il dimensionamento "Damage Tolerance" è lo strumento utilizzato per dimostrare che un danneggiamento prodotto per fatica o in modo accidentale nella struttura di un velivolo sarà rilevato, e conseguentemente riparato, prima che questo possa raggiungere una dimensione critica. Nel caso del nuovo addestratore M346, attualmente in fase di costruzione presso la AerMacchi SpA di Venegono Superiore, il dimensionamento Damage Tolerance viene applicato a tutti i componenti critici dal punto di vista della sicurezza del volo, nei quali un danneggiamento non rilevato causerebbe la perdita del velivolo. Da questo punto di vista uno degli elementi più critici è certamente l'attacco fra ala e fusoliera; per questo elemento è stato sviluppato, nell'ambito di una collaborazione fra il Dipartimento di Ingegneria Aerospaziale e la AerMacchi SpA, un programma di prove sperimentali a supporto della certificazione Damage Tolerance del nuovo addestratore. Nella configurazione attuale del velivolo i longheroni alari e le ordinate di forza, alle quali si collegano i longheroni, sono ricavate per fresatura da un massello prestirato in lega di alluminio 7050-T7451. Per la fabbricazione di questi elementi è stata adottata una tecnica innovativa di fresatura (High Speed Machining), nella quale si utilizzano velocità di rotazione dell'utensile molto elevate con piccole profondità di passata. Con questa tecnica si aumenta notevolmente la velocità di asportazione del materiale ed si abbassa il limite tecnologico sugli spessori minimi producibili imposto dalla fresatura tradizionale. Il collegamento fra longheroni e ordinate di forza è del tipo "lug-fork" con interposte, fra i perni in acciaio e la struttura in lega di alluminio, delle bussole speciali montate con forte interferenza, ForceMate®, anche al fine di aumentare la resistenza a fatica e rallentare la propagazione dei difetti. Il lavoro, tutt'ora in corso di svolgimento, ha riguardato la caratterizzazione, dal punto di vista della propagazione dei difetti, del materiale impiegato per la costruzione delle ordinate e dei longheroni. Essendo la propagazione influenzata anche dallo spessore del materiale, le prove sono state effettuate su provette CT aventi lo stesso spessore del lug, pari a 42 mm. Sono state eseguite prove sia con carico sinusoidale, sia con uno spettro di carico simulante le effettive condizioni operative dell'attacco, al fine di valutare gli "effetti di ritardo" caratteristici della sequenza di carico. Parallelamente è stato disegnato e costruito un provino di grandi dimensioni, rappresentativo dell'area critica a fatica nel collegamento ala-fusoliera; su questo provino 131 sono in corso prove di fatica tradizionali e prove di fatica in presenza di difetti iniziali introdotti per elettroerosione.

ATTIVITA' SU PROVINI "COMPACT TENSION"

Prove con carico sinusoidale. Questa attività ha riguardato la caratterizzazione del materiale utilizzato per la costruzione dei longheroni alari e delle ordinate di forza della fusoliera, lega leggera 7050-T7451. Ovviamente le proprietà delle leghe leggere sono reperibili nelle diverse banche dati esistenti, ma spesso non è possibile avere informazioni precise circa la sperimentazione che ha prodotto tali risultati. In particolare le proprietà di molti materiali tendono a migliorare nel tempo anche in virtù di miglioramenti nelle tecnologie di produzione. Per questo motivo si sono effettuate alcune prove di propagazione di difetti, con l'obiettivo di confrontare i risultati ottenuti con quelli disponibili nella banca dati di NASGRO, [1]. NASGRO è un programma, sviluppato dalla NASA, per la previsione della propagazione di difetti; tale programma è stato scelto dalla AerMacchi per effettuare le analisi Damage Tolerance. Poichè la velocità di propagazione di un difetto è notoriamente influenzata dallo spessore del materiale, le prove sono state effettuate su provette Compact Tension, CT, aventi uno spessore, 42 mm, uguale allo spessore del longherone nella zona di attacco ala-fusoliera; la dimensione in pianta del provino CT era W=240 mm, [2]. Con tali provette sono state effettuate tre prove nominalmente identiche per ciascuno dei tre diversi valori del rapporto di sollecitazione, R=Pmin/Pmax=0.02, 0.2 e 0.4. I risultati delle prove hanno mostrato delle caratteristiche di Damage Tolerance della lega 7050-T7451 più elevate rispetto a quelle contenute nella banca dati di NASGRO. Come esempio, la fig. 1 mostra i risultati ottenuti per le prove con R=0.02; la velocità di propagazione di un difetto è più bassa di quella prevista da NASGRO per tutti valori di DK (variazione nel ciclo di carico del fattore di intensità degli sforzi) ed il valore di soglia è più elevato. Una tendenza simile è stata osservata anche nelle prove con R=0.2 e R=0.4.

Resistenza a frattura. Secondo la banca dati di NASGRO, per il materiale in esame, spessore 42 mm, si ha un valore critico del fattore di intensità degli sforzi Kc=34 MPa m , mentre in alcune prove (R=0.02) si è osservata propagazione fino a valori DK=40 MPa m . Sono quindi state effettuate alcune prove W (mm) Kc ( MPa m ) per la misura della resistenza a frattura del materiale in esame 240 46.16 utilizzando due provini CT che si differenziavano per la dimensione 240 39.35 caratteristica: W=80 mm e W=240 mm. In entrambi i casi lo spessore 240 37.09 era B=42 mm. Le provette di dimensioni minori sono state costruite 80 34.44 in accordo con quanto previsto dall'ASTM [2], che consiglia, per le 80 34.81 prove di "fracture toughness" un rapporto W/B compreso 80 41.72 nell'intervallo 1¸2, mentre le provette di dimensioni maggiori erano Tab I - Risultati delle prove di al di fuori di questo intervallo; non esiste comunque un valido motivo fracture toughness per non accettare i risultati di prove effettuate su provette di dimensioni maggiori. I risultati ottenuti hanno mostrano una certa dispersione della resistenza a frattura, imputabile al materiale. Si vedrà in seguito l'influenza della resistenza a frattura anche sulle previsioni della vita a fatica. Dalle prove effettuate si può concludere che il valore di Kc contenuto nella banca dati di NASGRO, 34 MPa m è un valore realistico; nelle prove si sono comunque misurati valori fino a 46 MPa m .

Prove con spettro di carico. Sui provini CT (W=240 mm, B=42 mm) sono state effettuate anche prove di propagazione di difetti applicando una storia di carico rappresentativa della sollecitazione agente nel Lug, posto dalla parte del ventre alare, del longherone più sollecitato. La sequenza di carico è costituita da un blocco ripetitivo di 178 voli, che si differenziano in durata e severità. Il blocco ripetitivo equivale a 200 ore di volo; mediamente ciascun volo contiene 74.5 cicli di carico. La fig. 2 mostra l'andamento del carico agente sull'attacco durante i primi quattro voli della sequenza, adimensionalizzato con il carico massimo che si raggiunge una volta in 200 ore di volo. Scopo delle prove era quello di determinare i coefficienti del modello di Willenborg (generalizzato e modificato, MG), che fra le possibili opzioni di NASGRO, è il modello prescelto dalla AerMacchi per le analisi Damage Tolerance. Tale modello utilizza coefficienti di tipo semi-empirico, dipendenti dal materiale e dalla storia di carico applicata; la conoscenza di tali coefficienti consente l'applicazione del metodo a storie di carico sufficientemente "simili". 132

Nel modello di Willenborg generalizzato (WG), [1], si definisce un fattore di intensità degli sforzi "residuo", KR, tramite il quale si determina il rapporto di sollecitazione "effettivo", REFF, definito come: K - K K MIN R = MIN R = EFF EFF K - K K MAX R MAX EFF

Detta Z la dimensione della zona plastica associata ad un picco di carico, si ha: 2 p æ K ö Z = ç ÷ 8 è aS0.2 ø dove K è il fattore di intensità degli sforzi associato al picco di carico, S0.2 è la tensione di snervamento del materiale ed a è un coefficiente che dipende dallo stato di tensione all'apice del difetto (stato di tensione piano o stato di deformazione piano); valori tipici del coefficiente a sono compresi fra 1.15 e 2.55. Sia KOL il valore del fattore di intensità degli sforzi associato ad un picco di carico POL e ZOL la dimensione della zona plastica corrispondente; se per effetto dei cicli di carico che seguono il carico POL il difetto si è propagato di una quantità Da all'interno della zona plastica ZOL, si ha che il valore del fattore di intensità degli sforzi al quale è associata una dimensione di zona plastica uguale a:

Z= ZOL-Da è dato da: 2 æ Da ö K(*) = KOLç1- ÷ è KOL ø

Willenborg definisce fattore di intensità degli sforzi "residuo" la quantità: W (*) K R = K - K ac dove Kac è il fattore di intensità degli sforzi associato al valore corrente del picco di carico. Nel modello di Willenborg originale si ha ritardo se: W K R ³ 0 mentre nel modello di Willenborg generalizzato si utilizza un diverso valore del fattore di intensità degli sforzi residuo, dato da: W K R = f K R dove f è un parametro che dipende dal valore di soglia del fattore di intensità degli sforzi e dal rapporto OL K /Kac. In una successiva evoluzione del metodo, sviluppata per tener conto della diminuzione del ritardo che si ha nella propagazione dei un difetto in seguito alla applicazione di un carico di compressione (Modified Generalized Willenborg, MGW), e che costituisce il metodo utilizzato per il progetto Damage Tolerance del velivolo, il parametro f è dato dalla relazione:

2.523f f = o , R < 0.25 0.6 U 1+ 3.5(0.25 - RU ) f = 1 , R U ³ 0.25 dove RU è il rapporto di sollecitazione valutato sulla base del rapporto fra il carico di compressione e il sovraccarico di trazione. fo è il valore che il parametro f assume per RU=0; tale parametro dipende dal materiale e dallo spettro di carico e viene determinato sulla base dei risultati di prove di propagazione con spettro di carico. Valori tipici di tale coefficiente sono compresi fra 0.2 e 0.8. Le prove sui provini CT sono state effettuate con uno spettro di carico leggermente diverso da quello descritto in precedenza, ottenuto troncando tutti i carichi di compressione. La fig. 3 mostra i risultati delle tre prove effettuate con spettro di carico, confrontati, nella stessa figura, con risultati ottenuti con NASGRO per diversi valori del parametro di taratura del modello MGW, descritto precedentemente. E' evidente nella figura un diverso comportamento del modello rispetto ai risultati 133 sperimentali; il modello prevede la rottura finale con dimensioni di cricca di circa 100 mm, mentre sperimentalmente si sono osservate cricche aventi dimensioni fino a 130 mm. La rottura prematura prevista dal modello causa un forte aumento della velocità di propagazione anche con dimensioni di cricca relativamente piccola; variando il parametro di ritardo non si riesce a simulare il risultato degli esperimenti. Questo comportamento si giustifica con un valore non adeguato della resistenza a frattura del materiale; utilizzando nella simulazione il valore più elevato della resistenza a frattura misurato sperimentalmente, Kc=46 MPa m , si sono ottenuti i risultati mostrati in fig. 4. In questo caso si ha una buona simulazione dei risultati sperimentali; il valore più adeguato del parametro di calibrazione risulta essere: f0=0.62, mentre dai risultati riportati nella fig. 3 si sarebbe dedotto un valore f0=0.7, che è un valore molto elevato, anche se ancora nel range dei valori plausibili. Il modello di Willenborg è molto utilizzato per effettuare valutazioni sulla propagazione di difetti, ma rimane un modello di tipo semi-empirico che non tiene conto dei vari aspetti del fenomeno. Inoltre richiede una "calibrazione" e solo con una notevole esperienza è possibile stabilire i limiti entro i quali è possibile estendere i coefficienti determinati sperimentalmente. Più recenti, anche se ormai datati, sono i metodi basati sulla apertura del difetto, introdotta per la prima volta da Elber [3]; questi metodi tengono conto del fatto che il difetto rimane chiuso, e quindi non propaga, fino a che la tensione non supera un valore, Sop, che dipende dalla storia di carico trascorsa. Tali metodi si semplificano notevolmente se invece di calcolare la variazione della tensione di apertura durante la propagazione del difetto, si assume un valore costante per tale parametro, valutato sulla base del carico massimo e del carico minimo contenuti nella sequenza, [4]. Tale ipotesi è ragionevolmente applicabile a storie di carico brevi, per le quali la propagazione del difetto avviene all'interno della zona plastica prodotta dal picco di carico più elevato, fino alla presentazione dello stesso picco di carico nel blocco successivo. Per effettuare questa previsione si è utilizzato AFGROW, [5], un programma sviluppato da Air Force. AFGROW, fra le varie opzioni, ha un modello di propagazione con tensione di apertura costante i cui dettagli sono descritti in [6]. Per l'applicazione del modello è richiesta la conoscenza della tensione di apertura, Sop, misurata in una prova di propagazione con carico sinusoidale, R=0. Tale grandezza, che costituisce una proprietà del materiale, è stata misurata in apposite prove, non essendo reperibili in bibliografia informazioni relative a spessori elevati come quello di interesse. Dal punto di vista sperimentale esistono diverse tecniche per la misura della tensione di apertura, [7]. Nel caso in esame la misura è stata effettuata tramite un "clip gauge" montato sulla "bocca" del provino CT. Le rilevanti dimensioni del provino, W=240 mm, hanno notevolmente semplificato l'effettuazione della misura stessa. Nella fig. 5 è riportato un tipico risultato: la curva carico-apertura del difetto mostra una non linearità fino a che il difetto non è completamente aperto (la prova in questo caso è proseguita fino alla rottura del provino). Elber [3] aveva ipotizzato che il rapporto fra la tensione di apertura e la tensione massima dipendesse solo dal rapporto di sollecitazione, R. In realtà effettuando misure per diverse lunghezze di cricca si trova che tale rapporto non è costante, come mostrano i risultati ottenuti, fig. 6. Questo comportamento sembra essere una caratteristica di tutte quelle geometrie che, come il provino CT, presentano un valore elevato del gradiente dK/da, [8], che si traduce in un forte gradiente della dimensione della zona plastica nella scia della cricca. Non potendo introdurre in AFGROW un valore variabile del parametro Cf=Pop/Pmax, la simulazione è stata effettuata per piccoli incrementi della dimensione del difetto, in modo da poter assumere un valore costante di tale parametro all'interno di ciascun incremento. La fig. 7 mostra i risultati ottenuti; anche in questo caso la previsione è stata effettuata per due diversi valori della resistenza a frattura (34 e 46 MPa m ). I risultati ottenuti sono buoni, particolarmente utilizzando il valore più elevato di Kc (tale valore è anche il più plausibile, poichè la rottura di quasi tutti i provini è avvenuta nell'intorno di questo valore). E' interessante a questo punto rilevare una caratteristica del provino CT: in genere si afferma che il valore di Kc ha poca influenza sulla vita a fatica, ma il provino CT presenta un difetto iniziale grande, per cui la sua vita è relativamente breve rispetto a provini di geometria diversa. Si osserva quindi una dipendenza non trascurabile della vita a fatica dal valore critico del fattore di intensità degli sforzi.

ATTIVITA' SU PROVINI "ATTACCO ALA-FUSOLIERA"

Installazione delle bussole ForceMateÒ. La seconda parte dell'attività, tutt'ora in corso di svolgimento, ha riguardato un provino il cui disegno è stato ricavato dal disegno dell'attacco fra ala e fusoliera. In particolare il provino è un collegamento tipo "lug-fork" avente nella sezione critica la stessa geometria dell'attacco 134 inferiore del quinto longherone (il più sollecitato) alla rispettiva ordinata, fig. 8a. Le dimensioni principali del provino sono riportate nella fig. 8b; anche per questo provino il materiale è lega leggera 7050-T7451. Come di norma, il contatto diretto fra il perno e gli elementi strutturali viene impedito montando nei fori delle bussole con interferenza. La funzione delle bussole è quella di ridurre i problemi di fretting e di corrosione galvanica che sono tipici dei collegamenti lug-fork. Installando le bussole con forte interferenza si ottiene anche un effetto benefico dal punto di vista della resistenza a fatica, ma ciò è difficilmente realizzabile dal punto di vista pratico, anche per il rischio di produrre un danneggiamento durante l'installazione delle bussole. Il problema è stato risolto dalla Fatigue Technology Inc. (Seattle) con il sistema ForceMateÒ. Tale sistema prevede una bussola che viene montata con gioco nel foro, evitando quindi qualunque tipo di danneggiamento. Tramite una apposita attrezzatura, la bussola viene deformata radialmente, fig. 9, molto oltre il limite elastico del materiale costituente la bussola, in modo non solo da installare la bussola, ma anche produrre un cold working del foro. Questo è un ulteriore vantaggio dal punto di vista della resistenza a fatica. Su un provino è stata incollata una catena di 11 estensimetri per misurare le deformazioni prodotte durante l'installazione di una bussola e le successive deformazioni residue. Parallelamente è stata condotta un'analisi agli elementi finiti con il programma ABAQUS, [9], simulando numericamente l'installazione di una bussola. Per il calcolo erano disponibili le curve sforzo-deformazione dei materiali costituenti il provino e la bussola, 7050-T7451 e 17-4 PH, unitamente ai dati geometrici misurati prima dell'installazione, che indicavano un'interferenza del 2%. I risultati ottenuti sono mostrati in fig. 10, in termini di deformazione circonferenziale in funzione della distanza dal bordo del foro. A causa di un problema di saturazione del sistema di acquisizione dati non sono disponibili le registrazioni effettuate durante l'installazione della bussola. Confrontando le deformazioni lette con quelle calcolate dopo l'installazione della bussola si osserva un ottimo accordo, che induce a ritenere valide le deformazioni calcolate durante l'installazione e quindi le tensioni residue dopo l'installazione, rappresentate in fig. 11. Dal calcolo risulta una tensione di compressione nella bussola di circa 500 MPa; la zona interessata da tensioni residue di compressione si estende oltre la bussola per circa 5 mm, con un picco di compressione sull'elemento di alluminio di circa 180 MPa.

Difetti iniziali. In alcuni provini sono stati introdotti dei difetti iniziali per valutare il loro effetto sulla resistenza a fatica. In particolare, ovviamente su provini diversi, si sono introdotti difetti d'angolo e difetti centrali, come schematicamente rappresentato in fig. 12. Per i soli difetti d'angolo è stato possibile esaminare i due casi di difetto inserito prima dell'installazione della bussola e difetto inserito dopo l'installazione della bussola. Il difetto inserito prima dell'installazione della bussola è molto meno critico poichè l'espansione radiale prodotta durante l'installazione agisce come una sovra-sollecitazione che plasticizza tutta la zona circostante il difetto, creando ulteriori tensioni residue di compressione davanti al difetto stesso. Il difetto introdotto dopo l'istallazione della bussola, non usufruendo di questa sovra-sollecitazione, risulta essere più critico. I difetti iniziali hanno la forma di 1/4 oppure 1/2 circonferenza, raggio 1.27 mm e sono stati introdotti per elettroerosione.

Prove di fatica sui provini "Attacco ala-fusoliera". Allo stato attuale sono state effettuate cinque prove sui provini "attacco ala-fusoliera" di cui tre con bussole ForceMate e due comparative con bussole tradizionali (shrink fit, SF) montate con leggera interferenza, (0.05 mm), riscaldando la parte in alluminio e raffreddando la bussola con azoto liquido. I risultati di queste prove sono sintetizzati in fig. 12; l'attacco appare correttamente dimensionato a fatica, considerato che la vita di progetto è 15000 ore di volo. Questo obiettivo viene ampiamente superato anche con bussole SF in presenza di un difetto iniziale. Le bussole ForceMate forniscono un risultato decisamente superiore: due provini con difetto introdotto prima dell'installazione della bussola hanno superato le 60000 ore di volo simulate senza evidenza di propagazione. Un terzo provino con bussola ForceMate e difetto introdotto dopo l'installazione della bussola si è rotto dopo circa 65000 ore di volo. Non è possibile allo stato attuale effettuare una valutazione numerica della vita a fatica di un elemento con bussole ForceMate, per cui le attività future, oltre il completamento del programma di prove, riguarderanno questo aspetto; in particolare, con analisi agli elementi finiti, si intende valutare il fattore di intensità degli sforzi per un difetto presente in un lug nel quale è installata una bussola ForceMate. 135

CONCLUSIONI

Presso il Dipartimento di Ingegneria Aerospaziale dell'Università di Pisa è in corso di svolgimento un programma di prove sperimentali a supporto della certificazione Damage Tolerance del nuovo addestratore AerMacchi M346. In tale ambito, sono state effettuate prove di propagazione di difetti con spettro di carico su provini Compact Tension e prove di fatica/propagazione su provini di grandi dimensioni il cui disegno riproduce, in vera grandezza, l'area critica nell'attacco ala-fusoliera. I risultati delle prove hanno mostrato un comportamento del materiale utilizzato, lega leggera 7050- T7451, migliore di quello deducibile dalla banca dati di NASGRO, il programma con il quale vengono condotte dalla AerMacchi le analisi Damage Tolerance. Le prove con spettro di carico hanno consentito la valutazione del fattore di calibrazione del metodo Willenborg (generalizzato e modificato), necessario per tener conto degli effetti di ritardo della sequenza di carico. Sono state effettuate anche misure della tensione di apertura di difetti; i risultati ottenuti, utilizzati in un metodo di previsione semplificato, hanno fornito una buona stima della propagazione di difetti con carichi di ampiezza variabile. Nei provini attacco ala-fusoliera sono installate delle bussole ad alta interferenza, ForceMateÒ, la cui caratteristica è quella di aumentare la vita a fatica, oltre che risolvere i problemi di fretting e corrosione galvanica tipici dei collegamenti lug-fork. Sono state effettuate sia una misura che una valutazione numerica dello stato di tensione introdotto in un provino in seguito all'installazione della bussola. I risultati, in buon accordo fra loro, mostrano che durante l'installazione della bussola l'interferenza è tale da produrre anche il cold-working del foro. Ulteriori analisi numeriche riguarderanno il comportamento di difetti e la loro interazione con le autotensioni introdotte dalle bussole. L'attività sperimentale, ancora in corso di svolgimento, ha mostrato un buon comportamento dei provini anche con bussole tradizionali; la vita a fatica è risultata essere circa due volte la vita di progetto. Con bussole ForceMate, inserendo un difetto artificiale di 1.27 mm prima dell'installazione della bussola, non si è osservata propagazione dopo 60000 ore di volo simulate, pari a quattro volte la vita di progetto. Introducendo un difetto artificiale dopo l'installazione della bussola, condizione più gravosa che non la precedente, si è avuta rottura dopo circa 65000 ore di volo.

BIBLIOGRAFIA

[1] Anon. "Fatigue Crack Growth Computer Program Nasgro. Version 3.0: Johnson Space Center, JSC- 22267B, December 2000. [2] Anon. "Standard Test Method for Plane-Strain Fracture Toughness of Metallic Materials" in "Metal Test Methods and Analytical Procedures" ASTM-E399-90. [3] W. Elber "The Significance of Fatigue Crack Closure" Damage Tolerance in Aircraft Strucures, ASTM STP 486, pp. 230-247, 1971. [4] W. Elber "Equivalent Constant-Amplitude Concept for Crack Growth under Flight-Simulation Loading" Fatigue Crack Growth under Spectrum Loading, ASTM STP 595, pp. 236-250, 1976. [5] J.A. Harter "AFGROW User Guide and Thecnical Manual" AFRL-VA-WP-TR-1999-3016. 1999. [6] J.A. Harter "Comparison of Contemporary FCG Life Prediction Tools" International Journal of Fatigue, vol. 21 Suppl., pp. S181-S185, Nov. 1999. [7] W. Yisheng, J. Schijve "Fatigue Crack Closure Measurements on 2024-T3 Sheet Specimens" Fatigue and Fracture of Engineering Materials & Structures, vol. 18, n. 9, pp. 917-921, Sep. 1995. [8] J. Schijve "Fatigue Crack Closure: Observation and Technical Significance" Mechanics of Fatigue Crack Closure, ASTM STP 982, pp. 5-34, 1988. [9] ABAQUS Version 6.3-001, Copyrigth © Hibbitt, Karlsson & Sorensen, Italy. 136

10-2

7050-T7451; t=42 mm; R=0.02

10-3

10-4 Fig. 1 - Alcuni risultati delle prove di propagazione di difetti.

10-5 da/dn [mm/ciclo]

CT4 10-6 CT2 CT3 CT15 NASGRO

10-7 2 3 4 5 6 7 8 9 10 20 30 40 50

0.5 DK [MPa*m ] 1.0 Spettro LUG

0.8

0.6

0.4 P/Pmax 0.2 Fig. 2 - Primi quattro voli della sequenza di carico. 0.0 0 100 200 300 400 Tempo 500

-0.2 =0.61 =0.62 =0.63 =0.64 =0.65 =0.66 130 0 0 0 0 0 0 130 F F F F F F RISULTATI SPERIMENTALI 120 120

110 "MG" Willenborg =0.66 =0.67 =0.68 =0.69 =0.70 =0.71 =0.72 110 0 0 0 0 0 0 0 F F F F F F F 100 100

90 90 RISULTATI SPERIMENTALI "MG" Willenborg

80 80 a [mm] a (mm) 70 70

60 60 0.5 Kc=34 MPa*m0.5 Kc=46 MPa*m 50 50 Spettro CT. Pmax=91.8 KN Spettro CT. Pmax=91.8 KN 40 40 1 2 2 2 2 0.0 5.0x101 1.0x102 1.5x102 2.0x102 2.5x102 0.0 5.0x10 1.0x10 1.5x10 2.0x10 2.5x10 Ore (*200) Ore (*200) Fig. 3 - Risultati delle prove ad ampiezza variabile e Fig. 4 - Risultati delle prove ad ampiezza variabile e confronto con i risultati di NASGRO. confronto con i risultati di NASGRO. 137

100

80

Pop

60 Fig. 5 - Misura della tensione di opening.

P (KN) 40

CLIP 20 GAUGE

0 0.0 0.5 1.0 1.5 2.0 0.3 d (mm)

0.2 max /P op P 0.1 Fig. 6 - Tensione di opening in funzione della lunghezza del difetto.

0.0 0.2 0.3 0.4 0.5 a/W

130

120 AFGROW (Closure model)

110 0.5 0.5 Kc=34 MPa*m Kc=46 MPa*m Fig. 7 - Risultati delle prove ad ampiezza 100 variabile e confronto con i risultati di un metodo basato sulla tensione di apertura. 90

80 a (mm) 70 RISULTATI SPERIMENTALI 60

50 SPETTRO CT; Pmax=91.8 KN 40 0 10000 20000 30000 40000 50000 60000 Ore 138

Fig. 8a - Provino "attacco ala-fusoliera".

42

Fig. 8b - Provino "attacco ala-fusoliera"; dimensioni principali. R53 33 2 F

24.4

Fig. 9 - Installazione delle bussole "ForceMate".

Bussola ForceMate 139

12000

Deformazione durante l'installazione (FEM) Deformazione dopo l'installazione (FEM) 10000 Deformazione dopo l'installazione (estensimetri)

8000 y

) x 6000 me ( y e 4000

2000

0 0 2 4 6 8 10 12 14 16 18 x (mm) Fig. 10 - Deformazioni, misurate e calcolate, durante e dopo l'installazione di una bussola.

150

) 100 50 MPa (

0

y res -50 S -100 -150 -200 y -250 x -300 -350 -400 -450 -500 -550 -600 0 5 10 15 20 25 30 35 40 x (mm) Fig. 11 - Tensioni residue (calcolate) dopo l'installazione di una bussola. 140

7x104 corner crack (cc) c

6x104 a

5x104 surface crack (sc) 4x104

3x104 Ore di volo ForceMate + Difetto CC; a=c=1.27 mm Difetto SC + ForceMate; a=c=1.27 mm 2x104 Difetto CC + ForceMate; a=c=1.27 mm Nessun difetto Difetto CC; a=c=1.27 mm Vita di progetto 1x104

0 1 Shrink fit 2 3 ForceMate4 5

Fig. 12 - Primi risultati delle prove sui provini "attacco ala-fusoliera. 141 doi: 10.1111/j.1460-2695.2006.01062.x

Fatigue crack propagation in the wing to fuselage connection of the new trainer aircraft M346

A. LANCIOTTI1, F. NIGRO2 and C. POLESE1 1Department of Aerospace Engineering, University of Pisa, Via G. Caruso, 56122 Pisa, Italy; 2AerMacchi, a Finmeccanica Company, Via Foresio n. 1, 21040 Venegono Superiore, Varese, Italy

Received in final form 2 May 2006

ABSTRACT The series version of the M346 military trainer aircraft is currently under construction at Aermacchi (Venegono Superiore, Italy). The design target life of the aircraft, which will be certified for Damage Tolerance,is 12 000 flight hours (FH), with the possible extension to 16 000 FH after specific inspections. Fatigue tests were performed on critical elements at the Department of Aerospace Engineering at University of Pisa in order to verify crack propagation calculations. The wing to fuselage connection is one of the most interesting elements from the fatigue point of view. Spars and frames, both integrally machined, are connected by two lug-fork joints; the base material is aluminium alloy 7050-T7451 for both the elements. High interference bushings, ForceMate®, produced by FTI (Fatigue Technology Inc., Seattle, WA) were used in the lug/fork connections. Experimental activity was carried out on two different specimens. The first, a Compact Tensionspecimen, was tested under constant amplitude loading to verify the fatigue crack growth rate data contained in NASGRO 4, the software used for Damage Tolerance evaluations. Experimental results were fully comparable with the NASGRO 4 material database. Additional variable amplitude loading tests were carried out in order to calibrate crack growth prediction models used in the analyses. The second specimen was a lug-fork joint designed as the actual joints present on the aircraft. Both constant and variable amplitude loading fatigue tests were carried out in this case too. Results obtained clearly indicated the beneficial effect of ForceMate bushings, compared to shrink fit bushings.

Keywords ForceMate; lug-fork; variable amplitude loading.

possible extension to 16 000 FH after specific inspections. INTRODUCTION A two lifetime full scale test will be conducted to identify In a ‘Damage Tolerance’ design, a defect produced acci- critical locations; at the end of this test, artificial flaws will dentally or by fatigue in the structure of an aircraft, will be introduced in the main components and a third life- be detected and consequently repaired before it grows time full scale test will follow to verify damage tolerance to a critical dimension. A new trainer aircraft, M346, is characteristics of the airframe. currently under construction at AerMacchi (Venegono The wing to fuselage connection is one of the most crit- Superiore, Italy); all critical elements of this aircraft, i.e. ical elements. In the prototype configuration of the air- those that would cause loss of an aircraft in case of non- craft, each half-wing has three spars, connected to three detected damage, are designed on the basis of Damage main frames. Wing spars and frames are constructed by Tolerance criteria, following the guidance of U.S. Joint milling, using aluminium alloy 7050-T7451 plates in the Services Structural Guidelines, JSSG-2006.1 Design tar- prototype configuration, while 7050-T7452 plates will be get life of the aircraft is 12 000 flight hours (FH), with the used in the series configuration; such change is neces- sary because due to design modifications, thicker plates, available only for the second material, will be necessary. Correspondence: Agostino Lanciotti; Francesca Nigro, E-mail: a.lanciotti@ ing.unipi.it and [email protected] High speed machining, a technique based on high tool

1000 c 2006 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 29, 1000–1009 142 FATIGUE CRACK PROPAGATION 1001

revolution speed and high feed speed, is used to manufac- 10-2 ture these parts. High speed machining increases metal removing speed and, at the same time, allows machining 7050-T7451; LT; t=42 mm; R=0.02 of thinner parts compared with standard milling, thanks -3 to the lower forces exerted between the tool and the ma- 10 chined part. The connection between each spar and the correspond- ing frame is made by two ‘lug-fork’ joints; special bush- 10-4 ings, installed with high interference, named ForceMate®, da/dN [mm/cycle] produced by Fatigue Technology Inc., Seattle, are inter- posed between the steel pins and the aluminium structure. These elements are very efficient from the fatigue point 10-5 of view and in addition, suppress the fretting that typically occurs in lug-fork joints. Research activity was carried out in the framework of 10-6 CT4 (threshold test) a cooperation between AerMacchi and the Department CT2 of Aerospace Engineering at University of Pisa, to sup- CT3 port Damage Tolerance certification of the new trainer CT15 NASGRO 4 DataBase aircraft. Activity started with the characterization, from 10-7 the Damage Tolerancepoint of view, of the material used 2 3 4 5 6 7 8 9 10 20 30 40 in frame and spar construction of prototype aircrafts. Ad- ΔK [MPa m] ditional variable amplitude loading tests were carried out Fig. 1 Measured crack propagation rate in 7050-T7451 and in order to calibrate crack growth prediction models used comparison with NASGRO 4 database results. in the analyses. Subsequently, a larger specimen, with the same dimension as the most critical lower lug in the spar- frame connection, was designed and fatigue tested un- der constant and variable amplitude loading sequences, ments (42 mm) so as to take into account thickness effect representative of the aircraft current usage. The results on crack propagation rate; the characteristic dimension of this research activity, described in the present paper, of the CT specimen was W = 240 mm.3 Three identical will be useful also in the design of other aircraft critical crack propagation tests under constant amplitude loading elements. were carried out for three different values of the stress ratio, R = Pmin/Pmax = 0.02, 0.2 and 0.4. An additional threshold crack propagation test under decreasing load- DAMAGE TOLERANCE CHARACTERISATION ing was carried out in the case R = 0.02. OF THE MATERIAL USED IN THE WING TO Test results have shown that crack propagation proper- FUSELAGE CONNECTION ties of the material under examination compare well with information contained in NASGRO 4. As an example, Constant amplitude crack propagation and fracture Fig. 1 shows the results obtained at R = 0.02. A similar toughness tests trend was observed also in the tests with R = 0.2 and R = Crack propagation and fracture toughness tests were car- 0.4. ried out on the material under examination, aluminium In addition, six fracture toughness tests were carried out alloy 7050-T7451, to compare its properties with the on CT specimens having the same thickness (42 mm) and NASGRO material database.2 NASGRO is a software two different characteristic dimensions W = 84 and W developed at NASA for Damage Tolerance analysis; this = 240 mm. Results obtained were between 34.44 and . . program was chosen by AerMacchi to perform Damage 46.16 MPa m0 5; the mean value, 38.92 MPa m0 5, compare Tolerance evaluations on M346. well with the value contained in the NASGRO database, . In the wing to fuselage connection, each spar is con- equal to 35.16 MPa m0 5. Details of the test results are nected to the corresponding frame by two lug-fork joints; given in.4 lugs are in the spars while forks are in the frame. Spar lugs are more critical than fork lugs from the fatigue point Variable amplitude crack propagation tests of view, the acting stress being about 14% higher due to the difference in thickness. Constant amplitude crack Modified Generalized Willenborg (MGW) is, among dif- propagation tests were carried out on Compact Tension ferent options present in NASGRO, the model selected to specimens (CT), having the same thickness as the lug ele- evaluate fatigue crack propagation of the M346 structure.

c 2006 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 29, 1000–1009 1002 A. LANCIOTTI et al. 143

1.0 1.0

0.8 0.8 P/Pmax Δ 0.6

0.6 0.4 P/Pmax

0.4 0.2

0.2 0.0

-0.2 0.0 0.00.51.01.52.02.53.03.5 1 10 100 1000 10000 FLIGHT HOURS Number of exceedances in 200 flight hours Fig. 3 Loads acting in lower lugs of M 346 wing to fuselage Fig. 2 Non-dimensional load range exceedance diagram. connection during first four flights. =0.61 =0.62 =0.63 =0.64 =0.65 0 0 0 0 0

It uses an ‘effective’ stress intensity factor based on the φ φ φ φ φ size of the yield zone in front of the crack tip. A ‘tun- 120 ing factor’, 0, which depends on material and on the ‘kind’ of load sequence, is present in the model. This fac- 110 Experimental results tor, experimentally evaluated, can be used to predict crack CT 9 propagation in the same material under a ‘similar’ load se- 100 CT 8 quence. Typicalvalues of the tuning factor are in the range CT 11 0.2–0.8.2 90 Crack propagation tests were carried out on Compact 80 Tensionspecimens (W = 240 mm, B = 42 mm) to evaluate a [mm] the MGW model tuning factor, to be used for the load 70 sequence representative of the load acting in the most critical lower wing lug. This information would allow the 60 prediction of results of the tests subsequently performed on lug specimens. 50 Load sequence was composed of 178 flights of dif- CT Load Sequence. Pmax=91.8 KN ferent severity and duration; each flight contains on 40 0 10000 20000 30000 40000 50000 an average 74.5 load cycles. Whole sequence simulates FLIGHT HOURS 200 FH. Figure 2 shows a load range exceedance dia- gram, while, as an example, Fig. 3 shows the first four Fig. 4 Calibration of modified generalized Willenborg model. flights. The load sequence actually used in CT tests named ‘CT load sequence’, was slightly different com- pared to that described, which was used to test lug WING TO FUSELAGE CONNECTION SPECIMEN specimens. Compressive loads do not stress the critical The wing to fuselage connection specimen section of a lug joint, so compressive loads were trun- cated at zero level in CT load sequence; otherwise, com- Each M346 wing spar is connected by two lug-fork joints pressive stresses were maintained in fatigue tests of lug to a strong frame. The lower joint is the most critical one specimens, as they could negatively influence bushing from the fatigue point of view, being loaded by tension in installation. normal flight conditions. A specimen was designed start- Figure 4 shows results obtained; three identical tests were ing from the drawing of the actual structure as shown in carried out; result scatter was quite low. The MGW model Fig. 5; the specimen, Fig. 6, has the same dimensions as the calibration factor deduced from these tests is 0.625; nev- actual structure in the critical area. Two aluminium spac- ertheless, a conservative lower value of this parameter was ers, 1.4 mm thick, are interposed between lug and fork; used in the design of M346. these elements were introduced in the specimen design to

c 2006 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 29, 1000–1009 144 FATIGUE CRACK PROPAGATION 1003

Machined frame

Lug specimen Fig. 5 Specimen representative of the lower lug in M346 wing to fuselage connection. ForceMate bushing ForceMate

Fig. 7 Schematic representation of ForceMate bushing system. Fig. 6 Main dimensions of the specimen representative of the lower lug in M346 wing to fuselage connection. with slight clearance into the holes, to avoid possible in- take into account mounting tolerances between the three stallation damage to the hole; then bushing is radially ex- spars and the corresponding frames. A gap of 1.4 mm is the panded by a special mandrel, to produce an interference maximum tolerance admitted. Spacers were useful during fit mounting; the system is shown schematically in Fig. 7. tests because they were designed to allow the mounting Interference mounting is very high, so that fretting, typ- of crack propagation gauges in order to monitor lug crack ically observed in pin-loaded holes, is eliminated. The propagation. beneficial effect of interference fit on fatigue resistance Bushings are mounted both in lug and fork holes, just is well known, but in the case of the ForceMate bushing, as in the actual structure, to avoid direct contact between the hole is cold worked during installation, with an ad- aluminium components and 4340 steel pins. Special bush- ditional beneficial effect. A specimen was instrumented ings produced by Fatigue TechnologyInc., named Force- with strain gauges so as to quantify these effects; a chain Mate®,5 were used instead of common shrink fit bushings. gauge, composed of 10 measuring strain gauges with a ForceMate bushings, made of 17-4 PH steel, are mounted 1 mm pitch, was bonded on the specimen on the mandrel

c 2006 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 29, 1000–1009 1004 A. LANCIOTTI et al. 145 entry side (the left side of the work piece in Fig. 7), which 100 ) is free during bushing installation. The first strain gauge 50 MPa was located at 1.57 mm from the bushing edge. Bushing ( 0

was installed while strains were continuously recorded. res S -50

Subsequently, bushing was removed to put in evidence -100 the amount of residual strains due to bushing interference y -150 x and hole cold working. Bushing was removed by successive -200 reaming, to avoid damage to the specimen, while strains -250 were continuously recorded. The whole process was sim- -300 ulated by an ABAQUS finite element model;6 the mandrel was included in the model, together with the specimen and -350 the bushing. Cold working was realistically simulated by -400 ForceMate bushing installed -450 ForceMate bushing removed drawing the mandrel across the hole, from the entry face Shrink fit bushing (i=0.04 mm) to the exit face of the specimen, by imposing to it a transla- -500 tion displacement. Stress–strain curves of aluminium alloy -550 -202468101214161820 7050-T7451 for the specimen and steel 17-4 PH for the x (mm) bushing were introduced in the model, while the mandrel Bushing thickness was made of steel and had behaved elastically. Residual measured strains compared well with the numerical results Fig. 9 Residual stress with ForceMate bushing installed, residual predicted; as an example Fig. 8 shows residual strains af- stress after dismounting the bushing and residual stress due to ter bushing installation. Figure 9 shows the residual stress shrink fit bushing. field when bushing was installed and after subsequent re- moval. Bushing after installation is obviously compressed, improve fatigue resistance exploiting at the same time in- but also the material around the hole is compressed in terference and cold working. the circumferential direction, up to a distance of about A FEM model was used also to evaluate residual stresses 4 mm from the hole boundary. This is the consequence due to the installation of a shrink fit bushing having of the high interference during bushing expansion, about 0.04 mm interference; these bushings were used in com- 2%, which produces plastic strains around the hole. This parative fatigue tests. The stress field is completely differ- is, in all respects, a ‘cold working’ process; consequently, ent in this case; bushings produce tensile elastic stresses in after bushing removal, residual compressive stresses in- the lug (Fig. 9). By assuming that axial-symmetry exists, crease further, such as dimensions of the area subject to these stresses can be simply calculated also by the theory 7 compressive residual stress. In brief, ForceMate bushings of elasticity. The beneficial effect in this case and from the fatigue point of view8 is the reduction of fatigue stress ranges, even if residual stresses are low due to the small bushing interference adopted. 6000 FEM Experimental Initial defects Different initial defects were introduced in some speci- y mens to evaluate their effect on fatigue resistance. Defects 4000 x could be introduced before and after bushing installation. The second option was chosen, being the most critical; )

μγ indeed, plastic strains would have been produced around ( ε initial flaws during bushing installation. In this way initial flaws would have the benefit of an ‘over load’. A typical 2000 defect was a corner crack having a quarter circular shape, a0 = c0 = 1.27 mm, introduced by spark erosion after the installation of ForceMate bushings. Bigger initial flaws were also considered. 0 024681012Crack propagation under constant amplitude x (mm) loading on lug specimens

Fig. 8 Measured and evaluated residual strains after bushing Three comparative crack propagation tests under con- installation. stant amplitude loading were carried out on lug specimens;

c 2006 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 29, 1000–1009 146 FATIGUE CRACK PROPAGATION 1005 the first test specimen had a ForceMate bushing; the sec- bushing. It can be concluded that in ForceMate bushing ond test specimen had a shrink fit bushing installed with installation, the hole cold working effect prevails over the very small interference (0.02 mm) after having removed a bushing interference effect; indeed the beneficial effect ForceMate bushing. The third specimen was tested with- persists even if bushing is subsequently removed. out bushing; a 37-mm diameter pin was used in this case. Results obtained are of practical interest but the actual The aim of these tests was just to put in evidence sin- reason which motivated these tests was related to the be- gle factor contributions. Test conditions were: Pmax = haviour of the cracks observed in the tests carried out 245.25 kN, R = 0.05. Crack dimensions for the second under variable amplitude loading on specimens having specimen were available only by starting from a crack di- ForceMate bushings. In such tests, whose results are de- mension, a = 6 mm, so that all the results are compared by scribed later, it was observed that cracks grow faster in starting from this value. Results obtained, given in Fig. 10, the lug surface direction rather than the lug thickness show that, for the test conditions examined, ForceMate direction; this aspect, that could be put in relationship bushings increase crack propagation life about ten times; with residual stress field produced by ForceMate bush- results relevant to the specimen with the shrink fit bushing ings or variable amplitude loading, was observed also in installed with small interference after removing the Force- the specimen tested without any bushing under constant Mate bushing are in-between the other two cases, more amplitude loading, so that such reasons can be excluded. similar to results relevant to the specimen with ForceMate Coming to details, Fig. 11 shows the fracture surface of the specimen tested without bushing (Lug 14); an initial corner crack was introduced by spark erosion. Strangely, 30 three other significant fatigue cracks nucleated naturally P =245.25 kN; R=0.05 max during this test, while usually tests concluded with only 25 one large crack and some smaller fatigue cracks. Cracks grow with a ‘quarter elliptical shape’, maintaining the sur- 20 face dimension, ‘c’, larger than the depth dimension, ‘a’, while it is well known that the stress intensity factor in 15 a pin load joint containing a quarter circular shape crack

c (mm) is higher in the thickness direction compared with sur- 10 face direction; this result was confirmed experimentally in several papers, see for instance.9–13 The surface dimen- 5 Reference (No bushing) sion crack was monitored by a crack propagation gauge ForceMate ForceMate (removed) having twenty ties, pitch 1.15 mm, while crack depth di- 0 mension was numerically evaluated in such a way as to 0 100000 200000 300000 400000 500000 600000 transform the initial crack, having dimensions a0 = c0 = Cycles 2.3 mm, to the final crack observed in Fig. 11, having a = = Fig. 10 Crack propagation in lugs having shrink fit or ForceMate 13.4 mm and c 24.3 mm. These dimensions are com- bushings. pared in Fig. 12 with the results predicted by NASGRO

Fig. 11 Fracture surface of a lug specimen tested under sinusoidal loading.

c 2006 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 29, 1000–1009 1006 A. LANCIOTTI et al. 147

Fig. 12 Predicted and experimentally measured fatigue crack propagation in a lug specimen tested under constant amplitude loading.

and by AFGROW,14 a public domain fracture mechanic software for Damage Toleranceanalysis developed by the U.S. Air Force. This software contains a recent and ac- curate solution for the stress intensity factor of a corner crack in a pin-loaded hole.15 Results obtained show that a crack propagates very slowly compared with predictions, particularly in the thickness direction; AFGROW pre- dicts better the surface dimension of the crack, but in any case the result is very conservative. The crack shape is not Fig. 13 Crack observed in a specimen machined from the broken correctly predicted in either case. specimen shown in Fig. 9 (a), and from a different lot of the same Different hypotheses were formulated and checked by material (b). Finite Element analysis to explain the behaviour observed. Particularly, calculations were carried out to investigate the parameters not taken into account in Stress Intensity on specimens machined from different lots of the same Factor evaluation of pin loaded specimens, such as pin materials did not show this behaviour (Fig. 13b), so it elasticity and consequent bending with redistribution of is difficult to quantify the possible contribution of mate- contact pressure, friction coefficient in pin/hole contact rial anisotropy on the crack propagation rate in the short and clearance in pin/hole contact. Results obtained did transverse direction of lug specimens. Coming to the con- not give rise to correction factors for the Stress Intensity clusions of these tests, due to different and unidentified Factor which can justify the behaviour observed. effects, prediction of crack propagation in the specimen Suspecting that such behaviour was a consequence of ma- tested without bushing under sinusoidal loading was un- terial anisotropy, additional crack propagation tests were reliable; for this reason, for the time being, only a simple carried out on square section specimens, loaded in the description of the results obtained is given for the more rolling direction; these specimens were machined by using complex case of specimens having ForceMate bushings, broken lug specimens, so as to examine exactly the same tested under spectrum loading. material; material presented a ‘pancake’ microstructure with normal size grains surrounded by very small grains. Fatigue and crack propagation under variable This microstructure, considered normal in the production amplitude loading of these materials, promoted nucleation of cracks in the lug longitudinal direction, parallel to the material short Preliminary tests were carried out by applying the design transverse direction, as can be observed in Fig. 13a. This fatigue sequence; the maximum load in this sequence was phenomenon slows down the crack propagation rate in 426.5 kN. Testresults demonstrated that the specimen was the short transverse direction. Additional tests carried out over designed, so that the design life of 16 000 FH (with

c 2006 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 29, 1000–1009 148 FATIGUE CRACK PROPAGATION 1007

160000 Pmax=469.15 kN 140000 a=c= 0 mm 120000 mm a=c=1.27 a=c=4 mm

100000 a=c=1.27 mm

80000 mm a=c=6 a=c=6 mm 60000 a=c=6 mm FLIGHT HOURS FLIGHT

40000 16 a=c=2.3 mm

20000 mm a=c=2.3 Design life Defect inserted before FORCE MATE SHRINK FIT

13 15 6 7 91811 12 MATE/ FORCE 0 SHRINK FIT FORCE MATE

Fig. 14 Results of variable amplitude loading tests carried out on lug specimens. scatter factor 2) was widely exceeded. For this reason it had been decided to apply a 10% amplification factor to the load sequence; the design life was achieved in this case too. Figure 14 is a synthesis of all results relevant to lug speci- mens tested with the amplified variable amplitude loading sequence (maximum load 469.15 kN). All specimens were instrumented by crack propagation gauges, but compar- ison is here limited to total fatigue life, represented in Fig. 14 by a bar diagram. Small numbers in the lower left angle of each bar are specimen identification numbers which will be used in the discussion. Initial flaw dimen- sions are indicated over each bar. A vertical arrow indicates that a specimen was not completely broken at the end of Fig. 15 Fracture surface of a lug specimen tested under variable amplitude loading. the test. Specimen 9 was pristine, while different initial cracks were introduced in the other specimens by spark erosion. Initial cracks were corner flaws with the same di- ing bigger initial flaws, 4 and 6 mm, (specimens 18, 11 mension in the surface and thickness directions; flaws were and 12) good fatigue resistance was obtained. A 6 mm introduced after bushing installation, with the exception flaw was introduced in a specimen having a ForceMate of specimen 6 where the initial flaw was introduced before bushing removed after installation and substituted by a bushing installation. Two tests were carried out on speci- shrink fit bush. Fatigue resistance was good in this case mens having shrink fit bushings (i = 0.04 mm). The initial too, compared to ForceMate specimens having the same flaw was 2.3 mm in both tests; design life was exceeded in initial flaw (no. 11 and 12) due to the cold working ef- both cases, even if by a small margin. fect produced during ForceMate bushing installation. In The test on specimen 9, having a ForceMate bushing all tests, ForceMate bushings increased fatigue life more and no initial flaw, was suspended after 151 500 simulated than five times, compared to shrink fit bushings. FH, without evidence of fatigue damage. Even with the Severe fretting was observed in specimens having shrink introduction of a 1.27 mm flaw in specimen 7, fatigue fit bushings, while practically no fretting was observed in life was very high compared to design life. An initial flaw specimens having ForceMate bushings. As an example, introduced before bushing installation in specimen 6 did Fig. 15 shows the fracture surface of specimen 7, previ- not propagate during the fatigue test, so this condition ously described. The test was suspended at 115 800 FH, must be avoided, being less damaging. Even by introduc- when the crack was c = 24 mm. On the basis of results

c 2006 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 29, 1000–1009 1008 A. LANCIOTTI et al. 149 relevant to the other tests, it can be asserted that the was observed on some specimens extracted from the bro- residual fatigue life would be considerable, of the order ken lug specimens, but it was not observed in specimens of 30 000 FH. The specimen was partially saw cut, in or- obtained from different batches of the same material. der to easily break it and to observe the fracture plane. Variable amplitude loading tests were carried out by ap- No fretting was observed where the pin is housed. In this plying a 10% amplification factor to the load sequence, case too, the crack has grown more in the surface direction because preliminary tests demonstrated that the specimen than in the thickness direction. was over designed, so that the 16 000 FH design life was largely achieved in all conditions. In all tests ForceMate bushings increased fatigue life more than five times, com- CONCLUSIONS pared to shrink fit bushings. Crack propagation tests carried out on CT specimens The fatigue life of pristine specimens exceeded 10 times put in evidence fatigue crack growth rate properties of the design life. Even with the introduction of an initial the aluminium alloy 7050-T7451 fully comparable with flaw from 1.27 to 4 mm, the fatigue life was more than those contained in NASGRO 4, the software selected by five times the design life. Tests carried out on specimens AerMacchi for the Damage Tolerance design of the new having a shrink fit bushing installed with a slight interfer- M346 trainer aircraft. 7050-T7451 is used in the proto- ence after having removed a ForceMate bushing clearly type configuration of this aircraft to build, among other demonstrated that, from the fatigue point of view, the parts, wing spars and corresponding main frames by High cold working effect in a ForceMate bushing installation Speed Machining. prevails over the interference fit effect. A large lug-fork specimen, with the same dimensions Severe fretting was observed in specimens having shrink as the lower joints in the wing to fuselage connection, fit bushings, while practically no fretting was observed in was designed and fatigue tested under constant ampli- specimens having ForceMate bushings. tude loading and under the load sequence representative The activity carried out has clearly demonstrated that of the aircraft actual usage. Special high interference bush- the fork-lug joint was too over designed to support service ings, ForceMate, are mounted in these specimens. Strain loading. From a practical point of view, a consequence of gauges bonded on a specimen indicated that during bush- the results obtained is a possible reduction of the wing ing installation the hole is cold worked, with a further spars from three to two, after a complete revision of the beneficial effect from the fatigue point of view, besides aircraft design. interference fit mounting. In all tests a higher crack propagation rate was observed in the lug surface direction respect to the lug thickness REFERENCES direction, contrary to what has been observed in numer- 1 JSSG-2006 (1998) Joint service specification guide, aircraft ous papers from the literature. A test carried out under structures. Department of Defence. constant amplitude loading on a specimen not having 2 NASGRO 4.02 (2002) Fracture mechanics and fatigue crack a bushing installed demonstrated that crack shapes ob- growth analysis software. NASA Johnson Space Center and served were not in relation to variable amplitude loading Southwest Research Institute. 3 Anon. (1990) Standard test method for plane-strain fracture or residual stress field produced during bushing installa- toughness of metallic materials. Metal Test Methods and tion. Different hypotheses were formulated and checked Analytical Procedures ASTM-E399–90. by Finite Element analysis to explain the behaviour ob- 4 Lanciotti, A., Nigro, F., Polese, C. Damage tolerance served. Particularly, calculations were carried out to in- characterization of the wing to fuselage connection of the vestigate the parameters not taken into account in Stress aircraft M346. (In Italian). Paper presented at the XVIIth Intensity Factor evaluation of pin loaded specimens, such Conference of Italian Association for Aeronautics and as pin elasticity and consequent bending, friction coeffi- Astronautics, Roma, 15–19 September 2003. 5 Champoux, R. L. and Landy, M. (1986) Fatigue life cient in pin/hole contact and clearance in pin/hole contact. enhancement and high interference bushing installation using Results obtained did not give rise to correction factors for the ForceMate bushing installation technique. American Society the Stress Intensity Factor which can justify the behaviour Testing Materials, Special Technical Publication no. 927, pp. 39–52. observed. 6 ABAQUS Version 6.4, Copyrigth c Hibbitt, Karlsson & Optical microscope analysis of the material put in evi- Sorensen, Italy. dence a ‘pancake’ microstructure, characterized by nor- 7 Hsu, T. M. and Kathiresan, K. (1983) Analysis of cracks at an mal size grains surrounded by very small grains; this attachment lug having an interference fit bushing. American Society Testing Materials, Special Technical Publication 791, vol. 1, microstructure promoted nucleation of cracks in the lug pp. 172–193. longitudinal direction, parallel to the material short trans- 8 Anon (1985) Endurance of aluminium alloy lugs with steel verse direction, which slows down the crack propagation interference-fit pins or bushes. Engineering Sciences Data Unit, rate in the short transverse direction. This phenomenon no. 84025.

c 2006 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 29, 1000–1009 150 FATIGUE CRACK PROPAGATION 1009

9 Kathiresan, K., Brussat, T. R. and Rudd, J. L. (1985) Crack Computational and experimental fracture analysis of a growth analyses and correlations for attachment lugs. pin-loaded lug. In: BEASY Technical papers. http://www. J. Aircraft, 22, 818–824. cmbeasy.com/publications/papers/papers.html 10 Friedrich, S. and Schijve, J. (1983) Fatigue crack growth of 13 Kathiresan, K., Pearson, H. S. and Gilbert, G. J. (1983) Fatigue corner cracks in lug specimens. Delft University of Technology, crack growth of a corner crack in an attachment lug. Int. J. Report LR-375. Fract. 22, R59–R63. 11 Kim, J. H., Lee, S. B. and Hong, S. G. (2003) Fatigue crack 14 AFGROW Reference Manual (2003) Wright–Patterson Air growth behaviour of Al7050-T7451 attachment lugs under Force Base, AFRL/VASM. flight spectrum variation. Theor. Appl. Fract. Mech. 40, 15 Fawaz, S. A. and Andersson, B. (2004) Accurate stress intensity pp. 135–144. factor solutions for corner cracks at a hole. Eng. Fract. Mech. 71, 12 Curtin, T. J., Adey, R. A. and Brussat, T. R. (2006) 1235–1254.

c 2006 Blackwell Publishing Ltd. Fatigue Fract Engng Mater Struct 29, 1000–1009 SELECTED PAPERS I

4.2 INTEGRAL STRUCTURES BY INNOVATIVE JOINING TECHNOLOGIES

4.2.1 “FRICTION STIR WELDING” FOR AERONAUTICAL APPLICATIONS

[19] A. Lanciotti, C. Polese “Impact of Direct Water Cooling on the quality and performance of Friction Stir Welded joints in an AlMgSiCu aluminium alloy” Atti del Dipartimento di Ingegneria Aerospaziale, Università di Pisa, ADIA 2004-2, Settembre 2004...... 152

[26] T. Ghidini, C. Polese, C. Dalle Donne, A. Lanciotti “Simulations of fatigue crack propagation in Friction Stir Welds under flight loading conditions”, Materials Testing 7/8, pp. 370-375, 2006...... 172

151 152

DIPARTIMENTO DI INGEGNERIA AEROSPAZIALE

ATTI

A. Lanciotti, C. Polese

IMPACT OF DIRECT WATER COOLING ON THE QUALITY AND PERFORMANCE OF

FRICTION STIR WELDED JOINTS IN A ALMGSICU ALUMINIUM ALLOY

ADIA 2004-2

Settembre 2004

UNIVERSITA’ DI PISA 153

ABSTRACT

The paper describes the results of a test series carried on in order to characterise some welded joints realised on a 6013-T4 aluminium alloy with an innovative technology, “Friction Stir Welding”. The activity concerns some different aspects like static strength, fatigue resistance, thermal residual stresses and strains and fracture surface micro graphic analysis. Even if the welding process was not completely optimised, so that some defects being present in the joints, the results positively compared with tests previously carried out on the base material or on classic fusion welding joints. Friction Stir Welding seems to represent a valid alternative for joining big extruded aluminium panels in order to realise integral stiffened fuselage structures.

SOMMARIO

Si descrivono i risultati di una serie di prove di caratterizzazione effettuate su giunzioni in lega leggera 6013-T4 saldate con una tecnica innovativa, la “Friction Stir Welding”. Gli aspetti investigati hanno riguardato vari aspetti tra cui la resistenza statica, la resistenza a fatica, le tensioni e deformazioni residue di origine termica e analisi micrografiche delle superfici di frattura. Anche se il processo di saldatura non era perfettamente ottimizzato, essendo infatti presenti alcuni difetti nei giunti saldati, i risultati ottenuti si confrontano positivamente con i risultati di prove precedentemente effettuate sullo stesso materiale base e su giunti ottenuti con tecniche di saldatura per fusione classiche. La Friction Stir Welding potrebbe quindi costituire una valida alternativa per la giunzione di grandi estrusi in alluminio per la realizzazione di strutture irrigidite di fusoliera integrali.

154

1 INTRODUCTION

The need to design new commercial aircrafts to replace the existing ones at the end of their design life and to increase the world aircraft fleet that will grow in the next twenty years, has forced the aerospace International Community to investigate the possibility of obtaining more efficient structures in the cost and/or weight reduction field, with the same or better safety levels than the actual ones. Introducing new aircraft structures, cost and weight savings are pursued by an integrated design. The aerospace industry uses riveting as a fastening technology to join the aircraft panels of fuselage, wings and vertical and horizontal stabilizers. Nowadays great efforts are made to replace this standard joining methodology by new welding technologies which would provide considerable weight saving and cost reductions incorporating integral structures in aircraft constructions. Because of the combination of high strength and low weight, aluminium alloys are still very attractive for lightweight construction. However, certain aluminium alloys (2XXX and 7XXX series) are often proved difficult to weld by conventional fusion welding procedures. Friction Stir Welding (FSW), a new process developed by The Welding Institute (TWI, Cambridge, UK) in 1992 [1], is proving to be a viable joining process for all aluminium alloys and seems to be an effective process for replacing bolted and riveted joints. Typically, sheets and plates with a maximum thickness of 100 mm of 2XXX, 5XXX, 6XXX and 7XXX series aluminium alloys are being welded and big efforts have been made to understand the in- service behaviour of these systems. In a particular way, the structural component where attention has been focused is the fuselage, because a considerable reduction in weight and manufacturing costs are expected with the utilisation of this new joining technology.

1.1 THE PROCESS

Friction Stir Welding (FSW) is a solid phase joining technique developed by The Welding Institute (Cambridge, UK) and patented by TWI in 1991. Because of its solid phase, FSW allows the high quality joining of materials that have been traditionally troublesome to weld conventionally such as high-strength aerospace aluminium alloys like 2024 or 7475. Therefore, no undesirable, brittle, low-melting eutectic phases are formed during the welding process. The present status of FSW allows one to weld all aluminium alloys in thicknesses ranging from 1.2 to 100 mm, and even dissimilar aluminium alloys can be joined. A friction stir butt weld is produced by plunging a cylindrical rotating, non-consumable tool into the faying surfaces of the two plates to be joined (the work) [2]. The parts have to be clamped onto a backing bar in a manner that prevents the abutting joint faces from being forced apart. A drawing of a generic tool is shown in fig. 1: it consists of a shoulder and a profiled pin emerging from it. To initiate the welding process, the rotating tool is slowly plunged into the faying surface until some substantial portion of the shoulder is in intimate contact with the surface of the work-piece to avoid expelling softened material. In order to make a full penetration weld, the tool must be designed so that the bottom of the pin is near, but not at, the bottom surface of the work when the shoulder makes contact with the top surface. Hence, welding of various thickness work-pieces requires tools with different pin lengths or a special tool with a pin element of adaptable length or continuously retractable. Once the plunge is completed, travelling of the rotating tool along the joint line results in the formation of the weld, fig.1. In fact, by the friction and the pressure due to the relative motion, the metal edges warm and become plastic without reaching melting point. So the softened 155

weld metal is extruded around the pin and then forged by the trailing edge of the shoulder. The process can be visualised as an in-situ extrusion. The extrusion die travels along the length of the weld and is comprised of the backing plate or anvil (beneath the work-piece), the tool shoulder and pin, and the cold base metal of the work-piece. The longitudinal axis of the rotating tool is referred to as the forging axis and the travelling direction of the tool is the extrusion axis. The different existing shapes of the tool all promote high hydrostatic pressure along the joint line, causing consolidation of the plasticized material. The microstructure of a typical FSW joint will be briefly reviewed here. The extrusion and forging operations, coupled with tool rotation, are responsible for the characteristic microstructure that results from the FSW process as shown in fig. 2 with a micrograph obtained during the present activity, resulting from a FSW joint in a 6013 aluminium alloy. The weld microstructure features can be separated into two broad categories: the thermo-mechanically affected zone (TMAZ), and the heat affected zone (HAZ). The HAZ which will clearly lie closer to the weld bead, is similar to HAZs resulting from conventional, fusion welding processes: the material has experienced a thermal cycle which has modified the microstructure and/or the mechanical properties. Depending on the alloy, its initial heat treatment, and the proximity to the weld centreline, processes occurring in the FSW HAZ might include precipitate coarsening, precipitate dissolution, recovery, recrystallization, and grain growth. However, there is no plastic deformation occurring in this area. The TMAZ of a friction stir weld might be considered analogous to the fusion zone of a conventional weld except that, instead of being melted, the material in the TMAZ has been mechanically worked (plastically deformed by the friction stir welding tool). Within the TMAZ, there are three somewhat distinct regions. The first and most obvious is the “weld nugget.” The weld nugget is the region that has undergone the most severe plastic deformation and is characterised by a fine, relatively equiaxed and recrystallized grain structure. The width of the nugget is normally similar to, but slightly greater than, the diameter of the pin. Outside the nugget on either side is a second region, one that has been deformed to a lesser extent and that, depending on the alloy, may or may not show signs of recrystallization. In the case of aluminium the material can be extensively plastically deformed at high temperature without recrystallization and there is generally a distinct boundary between the recrystallized zone and the deformed zones of the TMAZ. However in most other materials the distinct recrystallized region (the nugget) is absent, and the whole of the TMAZ appears to be recrystallized. This is certainly true of materials which have no thermally induced phase transformation which will in itself induce recrystallization without strain, for example pure titanium, β titanium alloys, austenitic stainless steel and copper. In materials such as ferritic steel and α-β titanium alloys (e.g.Ti-6Al-4V), understanding the microstructure is made more difficult by the thermally induced phase transformation, and this can also make the HAZ/TMAZ boundary difficult to identify precisely. In fig. 2, the delineation between the nugget and the rest of the TMAZ is quite sharp because of the recrystallization resistance of the 6013 base metal. The deformation of the base metal grains manifests itself as bending in the plane of the metallographic section. The area immediately below the tool shoulder, which is clearly part of the TMAZ, should represent a separate category, as the grain structure is often different here. This third region (sub-zone) of the TMAZ is called the “flow arm.” The microstructure here is determined by rubbing by the rear face of the shoulder and the material may have cooled below its maximum. 156

It should be noted that the FSW process is not symmetric about the centreline: the advancing side of the weld is defined as the side on which the rotational velocity vector of the welding tool has the same direction as the travelling velocity vector of the tool relative to the work-piece. The retreating side is where the two vectors are of opposite direction.

1.2 PROCESS QUALITY

The difficulty of making high-strength, fatigue and fracture resistant welds in aerospace aluminium alloys (e.g. highly alloyed 2XXX and 7XXX series aluminium) has long inhibited the use of welding processes for joining aerospace structures. Instead extremely attractive results have been obtained when FSW has been used on high strength aluminium alloys. So Friction Stir Welding has recently been identified by leading aircraft manufacturers as “key technology” for fuselage and wing manufacturing [3, 4]. Besides this technology has been introduced into commercial practice in a number of applications [4-8], including also ship building and rail transport. Although many industries was looking at the process, enough information on the mechanical properties was not available to allow immediate applications. So big efforts were made by the international research community and industries too in order to prove that with this technology it is possible to obtain joints with high mechanical properties and excellent fatigue behaviour, comparable to the base material. There are a lot of process variables which can affect the quality of friction stir welds: tool design, tool rotational speed, travel speed, tool heel plunge depth, etc. and the base material as well. Although the vast majority of friction stir welds are expected to be flaw-free, it is not always possible to assume this. The principal aims are to fix welding procedures and material selection in order to enhance the confidence in the design and application of friction stir welded joints. The Department of Aerospace Engineering at the University of Pisa had developed a research activity on this argument to study in a deep and methodical way some fundamental aspects, in conjunction with the international industry aims [9]. The present study was performed in the context of a cooperation between the Department of Aerospace Engineering at the University of Pisa and EADS Corporate Research Centre Germany that realised the Friction Stir welded joints; the aim was to closely investigate the behaviour of FSW joints realised by a combined fluid-gas cooling device. The purpose was to verify the fatigue resistance of the FSW joints so obtained; attention was specially focused on the residual stress and strain distributions obtained with this new device.

2 EXPERIMENTAL ACTIVITY

The principal application for FSW joints is in fuselage and wing manufacturing. As shown in fig. 3 FSW technology can be intensively employed in fuselage welded integral structures e.g. in joining stringers or clips to the skin sheet or in butt joints of extruded panels with integral stringers. In the present activity the joining elements were extruded panels 3 mm thick with integral stringers; in correspondence with the weld bead the local thickness of the panels was increased to 4 mm in order to furnish sufficient material for the realisation of the bead as schematically shown in fig. 4. Two series of single sided butt welds, each approximately 700 157

mm in length, were produced; the welding direction was obviously parallel to the plate extrusion direction. Nowadays the aeronautical industry has been prompted to develop new weldable aluminium alloys, which, in addition to weldability, meet further requirements as aeronautic materials. The present investigation was performed using 3 mm thick Al-Mg-Si-Cu 6013 aluminium alloy, in the naturally aged condition T4. This material is used in automotive, railway and ship structures; besides it meets requirements as aeronautic material and is often presented as a possible alternative to the widely used AA2024-T3 Alclad in a corrosive environment. The main advantages of this alloy are a lower density (3%), a reduced fatigue crack growth rate, a higher yield stress (12-25%) and higher fracture toughness (23%); besides the higher elongation to failure often allows to obtain a formed sheet by only one mould. The base material was produced by OTTO FUCHS, Meinerzhagen, Germany. Although it had been proved that the Friction Stir Welding process is very robust in terms of the range of welding parameters that will produce acceptable welds, a large amount of testing and parameter tuning is necessary to create good results. In particular it is important to chose tool rotation and welding speed in order to minimise heat input which might introduce weakening effect in HAZ. In the present activity welding parameters had been selected by EADS Research Germany as a tool revolution of 900 min-1 and a welding speed of 800 mm/min, with an innovative designed and manufactured combined fluid-gas cooling device, fig. 5. The special behaviour of this cooling device is visible from the picture: the water cooling is working on the weld bead and the gas nozzle is “cleaning” the material in front of the weld seam in order to prevent excessive water at the joint line interface. It is the purpose of this system to minimise residual stress effects like distortion, possibly increase the corrosion resistance of welded joints and improve the overall strength properties of Friction Stir welded materials (smaller heat affected zone and enhanced artificial aging response). The tool employed was a 18 mm shoulder (D2 tool steel) with a M6 threaded pin profile. In fig. 2 is shown a typical macro- section of the weld bead so obtained. The objective of the present work was to characterise joint performance and surface features produced in the friction stir welds under “cooled welding conditions”. Two complete welded panels had been obtained under some slightly different conditions, giving rise to two different series of tests. The first one had been welded under the nominal condition parameters previous described; in the second panel the work butting central line had been intentionally laterally deviated with respect to the pin axis, in order to simulate a possible error in tool positioning in a large aeronautical structure, as the fuselage will be. The plates welded were subsequently sectioned transverse to the welding direction to obtain two series specimens. The specimen dimensions are shown in fig. 4. The weld face side of the weld bead was superficially finished by abrasives in order to remove superficial imperfections and achieve smooth joint surfaces.

2.1 ELONGATION TESTS

In order to provide global mechanical response of friction stir welding joints so obtained, some standard transverse tensile tests were conducted. Three standard, full-size transverse tensile specimens were machined from the first series welded plates so that the weld was centred in the gauge section and the loading axis was normal to the welding 158

direction; the global dimensions were the same as shown in fig. 4 but the central section was reduced from 38 mm to 14 mm in a dog-bone geometry. Tests were performed at room temperature on a 25-kN servo hydraulic test machine. Specimens had been instrumented by a 50 mm gauge length extensometer and loaded till final failure. The mean failure stress was 282 MPa; the elongation to failure was 3.8%. This datum requires some comments. All specimens failed in correspondence with the weld bead whose material has generally lower mechanical properties. The percentage elongation was calculated as if the entire gauge were deforming uniformly: the low value of the elongation to failure of FSW joints could give the wrong impression of a brittle fracture. Instead, when the welded joint fails, the base material is still in the elastic field. So it was important to measure the local elongation in the weld bead. Before load application, some marks had been introduced about every 2 mm along the specimens and the exact distance between every tick mark had been measured by means of an optical 100X stereo microscope equipped with an x-y table with digital display. Subsequently small load increments were applied by recording, at each step, the actual mark distances, the results obtained are given in fig. 6. The diagram shows an immediate onset of a non-uniform deformation, initially restricted to the region of the weld bead. The apparently lower elongation to fracture exhibited by the weld specimen is a result of the presence of a plastic strain limited to the weld bead but with a local peak value over 16%, fig. 6. Amazingly, in one test we obtained a local peak value above 25%. In this location however, some typical weld bead defects, described deeply later, were observed. Despite this, the maximum failure stress was about 75% of the base material final stress; this value is much more impressive if compared with the typical failure strength of joints realised with classical fusion welding technologies.

2.2 FATIGUE TESTS

Fatigue tests were carried out at room temperature by means of a 25 kN servo- hydraulic actuator. The test specimen geometry is shown in fig. 4; the welds were in the centre of the section and loading was transverse to the welding direction and the extrusion direction of the base panels. Stringers, ineffective in the test being perpendicular to load direction, were preliminary cut to facilitate the specimen machining. Tests were conducted at a nominal applied stress ratio of R=0.1, with different values of maximum stress, Smax and with a load frequency of 5 Hz.

2.2.1 RESULTS

Fig. 7 shows the fatigue results of all tests carried out at R=0.1 compared with the results of previous tests carried out on the 6013-T4 base material [10]. The scatter in the results was relatively low; the best–fit curve will be used to compared the results obtained. The high fatigue performance of FSW joints in 6013-T4 aluminium alloy is evident compared to the base material. The resistance of the welded joints is about 70-80% of the parent material resistance. This is a promising result from a design point of view, instead of the factor two or more in fatigue life obtained with traditional welding technologies. Results can be separately processed for the two series in order to show the effect of pin and butting line misalignment. Surprisingly results were not so different: instead tests relative to the first 159

series, with nominal parameters, showed a slightly higher scatter and a lower fatigue limit in the range of low stresses, fig. 8. On the contrary, important differences between the two welded panels clearly appeared by fatigue fracture surfaces analysis: each series showed a characteristic fracture surface feature, clearly correlated with its fatigue performance. So deeper microscopic analysis was carried out, as described later.

2.3 FRACTURE SURFACE ANALYSIS

Results obtained show the benefits of Friction Stir Welding for joining successfully also high-strength aluminium alloys. The low temperature of the so-called “solid state” welding process highly decreases the possibility of contamination of the metal. In practice it could be observed that some typical welding defects like porosity, shrinkage cracks and segregation disappear. However, these attributes do not mean that FSW material behaves like the base material. Although it has been proved that the FSW process is very robust in terms of the range of welding parameters that will produce acceptable welds, it has been observed that the weld bead can contain different types of defects such as discontinuities, micro voids and micro-structural imperfections. The presence of such discontinuities in the material can lead, during the in service life of the weld, to the start of cracks and, at the end, to the catastrophic failure of the structure. The second part of the experimental activity was concentrated on analysing the welding quality and the weld features obtained with this particular process. The location and shape of the defects could provide useful information concerning the development of the aluminium alloy friction stir welds.

2.4 EXAMINATION AT THE OPTICAL MICROSCOPE

At first, metallographic sections were used to try to detect flaws. The photos showed the characteristic joint microstructure with the banded structure of the nugget which forms the so-called “onion rings”: these are traces of the softened weld metal pattern when, during the process, it is extruded around the threatened pin; in any case, for the magnification employed, no voids, joint line remnants and root flaws appeared, fig.2. The fracture surface showed instead some of the characteristic defects of FSW technology. The surface marks clearly observed in both series were in the shape of chevrons, fig. 9; these features are the traces of the pin rotation i.e. the pitch between two successive signs is equal to the feed per revolution. The direction of the chevron marks is always opposite to the welding direction. Whether this is the cause of the features remains open to conjecture. “Tunnelling” could be clearly identified on the fracture surfaces of the first series specimen when microscopically examined, fig. 10. It could be observed also in the lateral surfaces of specimens as a visible void, fig. 11. However, the tunnelling was intermittent and it could not be identified in all specimen but in each case highlighted, the voids were on the retreating side of the weld and showed a clear delineation between the region of the nugget and the remainder of the TMAZ. Classic porosity in fusion welds is caused by gas bubbles, or by shrinkage at the time of solidification, and clearly neither of these processes is possible in friction stir welding. Such defects should therefore be referred to as “volumetric flaws” or 160

“voids”. When welding at high travel speeds, the material receives less work per unit of weld length, i.e. fewer tool rotations per unit length. Under such conditions, the plasticized material may be cooler, and less easily forged by the shoulder, resulting in voids remaining unconsolidated. As previously observed, this defect appears to be the most frequent cause of failure for the first specimen series. So it can be deduced that the cooled FSW process still needs a proper parameter arrangement to accommodate heat losses during processing. In fact clearly appears that samples tested in the present work had been welded with too low heat input parameters. Welding with this cooling device requires a higher tool revolution, about 1000/1200 rpm with an 800mm/min travel speed to overcome heat losses. The second series panel, obtained with the same parameters set up, shows instead a different failure mode. The introduced pin misalignment had been expected to give rise to some root defects, even if the external visual examination failed to detect the presence of root flaws. The visual and microscopic examination of the fracture surface revealed instead a line, slightly brighter in colour than the parent material, running the entire width of some specimens. The line was approximately 0.1-0.13 mm wide and positioned at the weld root, fig. 12. It does not necessarily represent an absence of any bond, indeed some regions of intermittent or weak bonding may be present. This type of defect represents the most common flaw observed in a single-sided Friction Stir weld; but flaw details will in turn depend on the welding conditions used to made the welds and the material that is being welded. So it has appeared useful to closely analyse this feature with proper methods, like bending tests.

2.5 BENDING TESTS

The difficulty in detecting root flaws using standard NDI methods is down to their nature, which had been shown to be one of intimate contact with not fully or poor strength bonding. Defects, not shown also by X-rays, may appear in the surface of a specimen when it is subjected to progressive localised overstressing. Several destructive tests were used to screen the welds for flaws, but only bend tests with the root in tension were found to be reliable: the opening of a crack in the weld root defined the weld as being flawed. In the present activity two bend tests for ductility of welds were conducted; root welds were positioned in the tensile side, according to the normative ASTM E 190 [11]. Specimens were milled in order to remove the local thicker material across the weld bead and obtain constant thickness. Welded joints did not pass the test: when specimens reached a bending angle of about 90° degrees, clearly visible flaws appeared on the root side, as shown in fig.13. Since the ASTM prescribes a flaw-free 180° degrees bending angle, the weld had to be defined as flawed. This destructive test has show the presence of the lack of penetration defect previously observed also in the fracture surfaces. The Friction Stir welds studied seem to be tolerant to the presence of root flaws without a significant performance penalty. Obviously the validity of this statement depends on the size and shape of the flaws actually produced. As Non Destructive Inspection methods of welds cannot be usefully employed to find root flaws, the process control during welding becomes a basic topic. Once the welding procedures have been defined, weld quality is maintained by control of the welding process and materials. Periodic destructive bend tests will enable ongoing quality to be monitored.

161

2.6 RESIDUAL STRESSES

Friction Stir Welding is seen as a promising method because of its limited heat input, compared with standard fusion welding techniques. Besides it had been proved that the low temperature of the so called “solid state” welding process highly reduces the residual stress distributions of the welding process in the joint (by definition, a distribution of stresses which are present in the material without the application of any external load). However, we have to consider that aluminium has a very high thermal conductivity which means that the heat produced by the friction of the tool spreads over a large area. Also the coefficient of thermal expansion is high and this, combined with the relative low yield strength and a rigid clamping, causes a certain amount of residual stress during the cooling of the weld area after FSW. The weld nugget and the heat affected zone after welding cool down, but the contraction of the material is restrained by the presence of rigid clamps and by the cooler aluminium parts. The result is the generation of longitudinal and transverse residual stresses [12]. The integrity of mechanical structures can be strongly influenced by the presence of residual stress distribution particularly when they experience in service a significant fatigue loading. As far as the present activity is concerned, one of the main goals of the cooling device employed was to strongly reduce the amount of geometrical defects and residual stresses in a welded joint. On visual analysis the welded panels showed a very limited amount of geometrical defects when compared also with others no cooled FSWelded or laser welded panels previously examined. Residual stresses were measured by using a destructive method: a panel was instrumented with strain gauges, as shown schematically in fig. 14. Strain gauges had been positioned in a section transverse to the weld bead on the tool side and oriented in the welding direction. Subsequently the panel was sectioned with a milling machine transverse to the welding direction in order to allow the relaxation of internal stresses. Attention was paid to protect strain gauges by sealant in order to avoid contact with hot metal shavings. Also the feed and rotational velocity had been selected in order to minimize plate heating. In this way the influence of the transversal stresses had been neglected, but usually these values are much lower than the longitudinal ones. Besides, the principal aim was a general evaluation of thermal residual stresses specially compared with those that compete with others welding technologies. In the case of the 6013-T4 aluminium alloy welded with this particular cooling device, a maximum value of 96 MPa was measured, slightly out of the weld zone, on the retreating side, fig. 15; the decrease of the residual stresses in the centre of the weld bead is the consequence of the lower mechanical properties of the welded material. This value is not as low as might be expected; for instance the residual stress obtained in a previous activity [13] with a 2219 alloy FSW gave rise to a lower peak value, 83 MPa, although the base material had higher mechanical properties and the tool was a “traditional” one. Anyway these values are considerably lower than the peak value measured on aluminium fusion welded joints; for instance the residual stress peak value measured in a 2219 [14] advanced Variable Polarity Plasma Arc welded plate was 140 MPa.

162

3. CONCLUSIONS

Different mechanical tests were carried out on 6013-T4 friction stir welded joints realised on integral extruded panels. The panels were provided by EADS Germany and were examined in the “as welded” condition even if better mechanical properties would be obtained after a T6 (artificial aging) thermal treatment. Compared with others FSW or laser welded panels previously examined, present ones showed a limited amount of geometrical defects thanks to the properly designed cooling device. The activity were performed on two large panels; the second one had been intentionally welded with a misalignment between the tool axis and the abutting line. The ultimate stress measured on welded joint was about 75% of base material reference value; elongation up to 25% were measured in the weld bead, while the base material was still in the elastic field. Friction stir welded joint fatigue performance compared quite well with the base material fatigue resistance; the fatigue results of specimens machined from the two different plates were fully comparable, even if different failure modes were observed. In particular lack of penetration was observed in the second plate and tunnelling was present also in the nominal parameter welded panel. This defect clearly demonstrated that samples had been welded with "too" low heat input parameters. So this technique will appear a promising one from a design point of view, but additional trials are necessary to eliminate these defects. A residual stress peak value of 96 MPa was measured in correspondence of the weld bead; this value is fully comparable with a value obtained in a classic FSW process. So the process does not appear so effective from the residual stress point of view even if the residual strains observed were very low.

ACKNOWLEDGEMENTS

The authors would like to thank F. Palm from EADS Germany, Munich, for kindly providing the FSW joints.

REFERENCES

1. W.M. Thomas, E.D. Nicholas, J.C. Needham, M.G. Murch, P. Temple-Smith, C.J. Dawes, Improvements Relating to Friction Welding, European Patent Specification 0 615 480 B1, 1992. 2. S. Kallee, D. Nicholas, Friction Stir Welding at the TWI, Copyright © 2004, TWI Ltd. 3. R.G. Pettit, J.J. Wang, C. Toh, Validated Feasibility Study of Integrally Stiffened Metallic Fuselage Panels for Reducing Manufacturing Costs, NASA/CR-2000-209342, 2000. 4. R. Talwar, D. Bolser, R. Lederich, J. Baumann, Friction Stir Welding of Airframe Structures, Proceeding of the 2nd International Symposium on Friction Stir Welding, Gothenburg, Sweden, 2000. 5. D. Waldron, Application of Friction Stir Welding for Delta Rocket Fuel Tanks, Proceeding of the 1st International Symposium on Friction Stir Welding, Thousand Oaks, California, USA, 1999. 6. I. Henderson, Exploiting Friction Stir Welding in Explosively Formed Aluminium Boat Hull Construction, roc. Conf. “INALCO” 98. 7th International Conference on Joints in Aluminium, Cambridge, UK, 1998, pp. 251-257. 163

7. T. Kawasaki, T. Makino, S. Todori, H. Takai, M. Ezumi, Y. Ina, Application of Friction Stir Welding to the Manufacturing of Next Generation “A-train” Type Rolling Stock, Proceeding of the 2nd International Symposium on Friction Stir Welding, Gothenburg, Sweden, 2000. 8. M. R. Johnsen, Friction Stir Welding Take Off at Boeing, Welding Journal 78 (2), 1999, pp. 35-39. 9. COFIN 2002, Programma di Ricerca Cofinanziato dal MIUR -Ministero dell’Istruzione, dell’Università e della Ricerca- “New Technologies for Pressurised Fuselages in Medium and Large Size Aircraft”, Coordinator G. Cavallini. 10. A. Lanciotti, Comportamento in Diversi Ambienti di Leghe di Alluminio di Impiego Aeronautico, Paper presented at the XIII Conference of Italian Association for Aeronautics and Astronautics, Roma, 11-15 September 1995 (in Italian). 11. Anon, Guided Bend Test for Ductility of Welds, American Society for Testing and Materials, E190-80. 12. C. Dalle Donne, E. Lima, J. Wegener, A. Pyzalla, T. Buslaps, Investigations on Residual Stresses in Friction Stir Welds, Proceeding of the 3rd International Symposium on Friction Stir Welding, Kobe, Japan, 2001. 13. A. Lanciotti, C. Polese, Some Properties of “Friction Stir Welded” Joints in two Aerospace Aluminium Alloys, Meeting “Friction Stir Welding for Aerospace Applications”, Pisa, 1 April 2003. 14. A. Lanciotti, A. Belmondo, Mechanical Properties of Al 2219 VPPA Weldments, 48th International Astronautical Congress, Turin, Paper IAF-97-1.04.07, 1997.

164

Fig. 1 - Friction Stir Welding process.

Fig. 2 - Metallographic section of a AA6013-T4 weld.

165

Frame Today Clip Differential Structures

Stringer T-joint (Clip/Sheet) Skin Sheet

Tomorrow

Integral Fuselage Design

T-joint (Stringer/Sheet) Butt joint (Sheet/Sheet)

Fig. 3 - Future application of FSW technology.

38 30

4 3

260

Fig.4 - Fatigue test specimen, dimension [mm]. 166

Fig. 5 - Cooling device.

Fig.6 - Elongation tests. 167

200 180

160 6013 T4 base material 140

120

,[MPa] 100 max S 6013 T4 FSW 80 → → → → 60 →

10000 100000 1000000 1E7

Fig.7 - Fatigue test results.

200 180

160 Series 2 140

120

100 Series 1 S max, [MPa] S max, 80 → → 60 →

10000 100000 1000000 1E7 Cycles N Fig.8 - Tests results. 168

CHEVRON

Fig.9 - Chevron features on the fracture surface.

TUNNEL

Fig.10 - Tunnelling on the fracture surface. 169

0.454 mm

Fig.11 - Tunnelling on the lateral surface.

LACK OF PENETRATION

0.113 mm

Fig.12 - Lack of penetration on the fracture surface.

170

Fig.13 - Root bending test.

Fig. 14 - Residual stresses measurement. 171

100 6013-T4 FSW 2219-T851 FSW 80

60

40

20 Retreating side Advancing side FSWeld 0

-20

-40 Longitudinal residual stress [MPa]

-100 -50 0 50 100

Di t f th t li f th ld [ ] Fig.15.- Residual stresses.

MP 100746 – Text geliefert – 1 Tabelle gesetzt

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Simulations of Fatigue Crack Propagation in Friction Stir Welds under Flight Loading Conditions*

Tommaso Ghidini, Claudia Polese, Since Friction Stir Welding has been identified as a key technology Bremen, Agostino Panciotti, Pisa, for primary aerospace structures, fatigue crack growth testing and Italy and Claudio Dalle Donne, modelling is required. However, there is still a complete lack of Munich information regarding the fatigue crack propagation in friction stir welds under variable amplitude loading and flight loading condi- tions. The experimental investigations have been performed on cen- tre cracked 4 mm thick AA2024-T3 base metal and FSW specimens: regarding the FSW samples, the crack was placed 5 mm out of the weld centre, in the most critical part of the joint: All the coupons have been tested under simple variable amplitude load sequences, and under a standardized flight-simulation load history, FALSTAFF. The fatigue crack propagation was then predicted using wide- spread aerospace fracture mechanics software packages. Inter- action effects and internal residual stresses were firstly separately simulated and than combined in order to evaluate the ability to predict the fatigue crack propagation on FSW welded structures under service loading conditions.

Friction Stir Welding [1], FSW, is essen- this specific joint type: the life predic- quantitative predictions is still a formi- tially a solid state welding technique tions are the basis for scheduling period- dable task. The literature concerning fa- which allows high quality welds without ic inspections and for the maintenance tigue crack growth under variable am- solidification cracking, porosity, oxida- programs, which are costly and time plitude loading is voluminous and a tion and other defects typical to tradi- consuming. From a safety point of view number of excellent review papers are tional fusion welding. In aeronautics, it is important that they are conservative also available [5], [6]. However, there is friction stir welding is considered as key enough to prevent a significant loss of a complete lack of information regarding technology to reduce aircraft assembly structural integrity from cracks occur- the fatigue crack propagation in friction cycle time, cost and weight while main- ring between inspections, but from an stir welds under variable amplitude taining good strength and fatigue prop- economic perspective they should not be loading and flight loading conditions, erties [2] – [4]. The recent FAR regula- too conservative. both from an experimental as well as tions for damage tolerance and fatigue The loading experienced by aircraft from a prediction point of view. evaluations of aircraft structure require structures is generally variable ampli- Referring to the crack propagation on fatigue crack growth predictions. If fric- tude loading: the different types of load a welded structure, his integrity can be tion stir welding is considered for pri- sequences are known to induce a num- strongly influenced by the presence of mary aircraft structures, the prediction ber of different load interaction effects, residual stress distribution particularly methodologies have to be adopted for retardation and acceleration, which can when they experience in service a sig- result in significant variation on the nificant fatigue loading [7]. The very crack growth rate. Unfortunately, there rigid clamping arrangement used dur- * Work performed during the staying of the au- are so many interaction effects between ing FSW exerts a high restraint on the thors at the Institute of Materials Research, load cycles of different amplitudes that, deformation of the welded plates: these German Aerospace Center, Cologne Ger- many. This contribution was already pub- even if some of them are quite well un- restraint impedes the contraction due to lished in German in the DVM-Report 132 – derstood (at least from a phenomenolog- cooling of the weld nugget and heat af- Joining and Fatigue ical point of view), carrying out accurate fected zone in longitudinal as well as in

2 © Carl Hanser Verlag, München MP Materialprüfung Jahrg. 48 (2006) 7-8 173 BETRIEBSFESTIGKEIT MP

transverse direction, introducing trans- verse and longitudinal residual stresses [8] – [11]. In the present investigation an exploratory test program has been car- ried out to improve the general under- standing of the sequence effects in FSW Figure 1. M(T) fatigue crack propaga- tion test specimen. The detail in the box joints. The aim is to offer a basis both for shows the crack starter, the sketched a better experimental documentation as specimen is the FSW specimen, same well as for an evaluation of existing coupons were machined for the base crack propagation prediction models. material Moreover specifically existent fracture mechanics principles are applied to fric- tion stir welded structures, in the pres- ence of residual stresses. The objective is to elaborate an easy, practical and re- liable engineering approach to perform life predictions of friction stir welded Fatigue Crack Propagation Testing. Fatigue crack propagation under vari- structures under in-service loading. Variable amplitude loading testing was able amplitude loading as well as under performed on the AA 2024-T3. All the flight simulation loading were carried Experimental Program coupons were tested under simple vari- out on a servo-hydraulic machine of 400 able amplitude load sequences such as kN capability controlled by the Wave Friction Stir Welding and Specimen single overloads, overloads and under- Maker software (for the variable ampli- Geometry. The material used on the loads, overloads and compressive under- tude loading tests) and by the Fast Track entire experimental program is a 4mm loads, and under a standardized flight- software (in the case of the flight simula- thick AA 2024 in T3 condition provid simulation load history , FALSTAFF [12]. tion tests). ed by Pechiney in the form of sheets. In Figure 2 are specified the different Regarding the variable amplitude The 4 mm thick aluminium alloy sheets variable amplitude load types which has loading testing, all the sequences have were welded on a conventional milling been used for the experimental pro- been applied iteratively till the speci- machine at DLR, the German Aerospace gram. The load values are all in MPa and men’s failure. The fatigue crack propa- Centre in Cologne. All welds were butt the spectrum types have been derived gation rate has been measured by using joints and they were performed par- from a previous work [13] in order to the potential drop method [14]. To allow allel to the rolling direction of the profit from the gained experience and compressive loads during the tests the sheets. for reference purpose. FALSTAFF is a specimen were enclosed in anti-buck- The material was welded in T3 condi- standard load sequence considered rep- ling guides. The tests were carried out in tion and no additional heat treatment resentative for the load time history in a laboratory air environment at room was carried out after the welding proce- the lower wing skin near the wing root temperature, all in duplicate. Concern- dure. The welding speed was 300 mm of a fighter aircraft. The FALSTAFF with ing the flight simulation testing, the fa- per minute, the rotating speed of the tool a maximum of 196 MPa has shown the tigue crack propagation rate was mea- was 850 revolutions per minute and the more repeatable fatigue crack propaga- sured optically by stopping the test tilt angle of the tool was kept constant at tion life and therefore was applied also every 200 flights. The experimental re- 0°. The friction stir welded fatigue speci- to the friction stir welded specimens. sults of the base and FSW specimens mens were machined in the T-L orienta- tion (the loading axis was perpendicular to the rolling direction and to the weld- Load type ing direction) with the weld nugget in MPa the middle of the coupon: a crack starter 80 Constant amplitude notch consisting of a 4 mm diameter 1 8 R = 0.1 hole and two saw cuts (width 3 mm) with a total nominal length of 10 mm was ma- 100 Spectrum 1 chined in the centre of the base material 2 60 Variable amplitude 20 Overload specimens. In the FSW coupons it was 5 100 Figure 2. Load types mm out of the weld centre (Figure 1) 140 Spectrum 2 of constant and variable which is the most critical part of the 100 Variable amplitude amplitude load histories 3 60 Overload-underload joint: the crack was placed in the Ther- 39 20 momechanically Affected Zone (TMAZ) 140 Spectrum 3 (the grains are plastically deformed by 100 Variable amplitude 4 60 Overload-underload the tool and a partial re-crystallisation is -37 39 (compression) present due to the heat input from the shoulder), on the retreating side, where 5 Falstaff the minimum value of the hardness is lo- cated [8].

Jahrg. 48 (2006) 7-8 3 MP BETRIEBSFESTIGKEIT 174

20 [19] and Nasgro 3.0 [20]. The codes were MPa A y already successfully used from the au- 10 thors for predicting fatigue life and fa- res R x σ tigue crack propagation of base as well 5 Figure 3. Transversal as friction stir welded specimens [21, ress residual stresses distri- 0 22]. bution as a function of -5 As previously underlined the most ev- the distance from the -10 ident effects acting on a structure sub-

weld centre [18] Residual st jected to variable loading histories are -15 Cut compliance [20] the retardation and acceleration of the -20 0 10 20 30 40 50 60 mm 80 fatigue crack growth, the so called inter- x action effects. In welded structures, the fatigue crack propagation is affected by have been plotted in a-N diagrams, the welding direction in order to allow the presence of residual stresses. Those where N is now the number of flights. the relaxation of internal stresses. Atten- two macroscopic effects will be firstly Residual Stresses Measurement.The tion was paid to protect strain gauges by separately simulated and than com- determination of the residual stresses sealant in order to avoid contact with hot bined. To keep the model simple and re- was obtained with the so called cut com- metal shavings. In this experimental test liable the following basic assumptions pliance method [15–17]. The cut compli- the cut was performed on the centre of are pursued: ance method is a destructive technique the specimen moving in both direction to • base material dadN/ΔK curves will be which directly gives the residual stress- the edges of the coupon: the strain used for all the predictions (also for es or the residual stress intensity factors gauges have been placed 5 mm out of the the welds) simply performing a narrow saw cut, in- weld in order to measure the residual • existent retardation models with sug- troduced progressively every millime- stresses where the crack is moving. The gested parameters will be used to pre- tre, on the tested specimen. Every mil- strain gauges were applied parallel and dict the fatigue crack propagation un- limetre of cut a certain amount of resid- transversal to the direction of the cut: der variable amplitude loading ual stresses is released: the specimen this was done because the measuring re- • and an easy approach to measure the residual stresses will be adopted in 103 the case where the residual stresses mm/cycle AA2024-T3 NASGRO 3.0 Database are not negligible and they will be Experimental data 101 simply added to the base material 100 da/dN-ΔK curves. Figure 4. Comparison 10-1 A same approach was successfully -2 of the experimental 10 adopted in the prediction of an easiest fatigue crack growth 10-3

da/dN spectrum on FSW CT specimen [22]. In curve and the material -4 10 modeling the fatigue crack propagation stored in the NASGRO 10-5 tests the base material data stored in the database 10-6 10-7 Nasgro database, which was processed -8 by the so called Nasgro equation [23] 10 —— 1 10MPa√m 100 was used. The material constants are DK those for the AA 2024-T3 orientation T(L), as listed in Table 1. In Figure 4 the reacts to this redistribution of residual gards the transversal residual stresses: curve obtained with these constants and stresses with a deformation which is since the longitudinal residual stresses adopted for the calculations and the measured as strain by a strain gauge are bigger than the transversal they have experimentally measured crack growth glued on the specimen. also to be measured. The residual stress curve of the base material , are com- A panel of the same dimensions of the distribution obtained by this simple pared in da/dN-ΔK diagram. The ma- fatigue crack propagation specimens method is reported in Figure 3. terial tested has a faster crack growth was instrumented with strain gauges rate than the material used in the pre- [18]. Strain gauges were positioned in a Fatigue Crack Growth diction. section parallel to the weld bead on the Predictions The dimensions of each experimental tool side and oriented in the welding di- specimen have been used in Afgrow and rection. Subsequently the panel was sec- The calculations have been done with in Nasgro 3.0 with the proper initial tioned with a milling machine parallel to two fracture mechanics software Afgrow crack dimensions for the implemented model called Internal through crack, in Afgrow, and Center cracked tension in 3/2 Δ 3/2 α UTS [MPa] YS [MPa] KIC [N/mm ] K0 [N/mm ]C npq Nasgro. 448.16 330.95 1007.7 100.77 6.08x10-11 2.6 0.5 1.0 1.5 The model implemented in Nasgro uti- lizes a stress intensity factor solution Table 1. Material parameters given in the material database of NASGRO for the material developed [24, 25] while the one used in 2024-T3 T(L) Afgrow refers to [26].

4 Jahrg. 48 (2006) 7-8 175 BETRIEBSFESTIGKEIT MP

The implemented retardation models of the fracture mechanics software [27 to 29] were firstly tested on the base mate- rial with the easiest spectrum loading of Figure 5. Overview of the test program: the spectrum 1 of the complete experi- Fehler! Verweisquelle konnte nicht mental results: 2024-T3 base material and FSW gefunden werden. The retardation mod- Bild 5 fehlt! experimental observa- el giving the best results was then used tions are compared for to predict the other crack propagation every single loading experiments under more complex load- spectrum ing histories. Regarding the welded specimens, only Afgrow was used for the calculations due to his capability to take into account the residual stresses by meaning of implemented weight func- developed: it may provide a practical two different problems have been then tions. After calculating the residual and reliable basis for the analysis of fa- superimposed to predict the fatigue stress intensity factor, Kres, the code tigue tests of integral structures under crack propagation of friction stir welded simply add it to the stress intensity in-service loading conditions, by using specimens under more complex loading factor due to the external load, as sug- widespread fracture mechanics soft- spectra and under flight simulation. The gested by the superposition principle ware. The fatigue crack propagation of simple measurement of the residual [30]. friction stir welded specimens has been stresses and the inclusion of the retarda- predicted not only under constant ampli- tion models of the software packages im- Results and Discussions tude loading but also under variable am- proved the quality of the predictions plitude loading histories and under considerably. Particularly for loading Figure 5 shows the complete program flight simulation. The two most impor- spectra with a predominant influence of results of the AA 2024-T3 as comparison tant effects affecting the fatigue crack overloads the Willenborg model with a between the base and the friction stir propagation under complex and in-ser- shutoff overload ratio RSO = 3 gives the welded material for every single loading vice loading conditions of welded struc- best results, while for loading spectra spectrum summarized in Figure 2 in tures have been identified and firstly with a predominant influence of under- crack length versus number of cycles (or separately simulated: the residual loads the best predictions have been ob- number of flights in the case of the FAL- stresses and the retardation and acceler- tained without using any retardation STAFF spectrum) diagrams. All the com- ation effects. The single solutions of the model. parisons of Fehler! Verweisquelle konnte nicht gefunden werden. refer 6 to one base metal curve and two FSW 10 Base Material curves in order to have information also Spectrum 1 about the scatter in the welded struc- Spectrum 3 tures. As general trend the FSW speci- 105 Constant amplitude Spectrum 2 mens showed a faster fatigue crack Non-conservative Figure 6. Predicted growth rate as was expected by consid- Predicted atigue life of the base ering the presence of tensional residual 4 material compared

Cycles 10 stresses presented in Figure 3. An over- Conservative to the actual one all comparison between experimental Falstaff (Number of flights) and predicted values of base and friction 103 stir welded 2024-T3 is presented in Fig- 103 104 105 106 ure 6 and Figure 7. CyclesExperimental The estimated lives are in good agree- ment with the actual lives showing the viability of the assumptions made in the 106 analysis. The observed non-conservative FSW Material predictions could also be explained by considering the different crack growth Spectrum 3 Spectrum 1 105 rate used by the codes in performing the Constant amplitude Spectrum 2 Figure 7. Predicted calculations. Predicted Non-conservative fatigue life of the friction 4 stir material compared

Cycles 10 Conclusions to the actual one Conservative A simple engineering approach which is Falstaff (Number of flights) based on a relatively solid background 103 3 4 5 6 and which is checked against fatigue 10 10 10 10 test data for various test conditions was CyclesExperimental

Jahrg. 48 (2006) 7-8 5 MP BETRIEBSFESTIGKEIT 176

References 14 Bachmann, V.; Marci, G.; Sengenbusch, P.: 22 Ghidini, T.; Dalle Donne, C.: Prediction of Procedure for Automated Tests of Fatigue Fatigue Crack Propagation in Friction Stir 1 Thomas W. M.: Friction Stir Butt Welding, Crack Propagation, ASTM STP 1231 (1992), Welds, 4th International Conference on International Patent Application pp. 146-163 Friction Stir Welding, Park City, Utah (2003) No. 9125978.8 (1991) 15 Cheng, W.; Finnie, I.: Measurement of Resi- 23 Forman, R. G.; Mettu, S. R.: Behavior of 2 Christner, B.; McCoury, J.; Higgins, S.: dual Hoop Stresses in Cylinders Using the Surface and Corner Cracks Subjected to Development and Testing of Friction Stir Compliance Method, Journal of Enginee- Tensile and Bending Loads in Ti-6-Al-4V Welding (FSW) as a Joining Method for Pri- ring Materials Tech. 108 (1986), pp. 87-92 Alloy, Fracture Mechanics 22th Sympo- mary Aircraft Structure, 4th International 16 Schindler, H.-J.; Finnie, I.: Determination sium, ASTM STP 1131 (1992) Symposium on Friction Stir Welding, Park of Residual Stresses and the Resulting 24 Feddersen, C. E.: Discussion of Plane Strain City, Utah (2003) Stress Intensity Factors in the Ligament of Crack Toughness Testing of High Strength 3 Sheperd, G. E.: The Evaluation of Friction Pre-Cracked Components, Proceedings Metallic Materials. ASTM STP 410 (1966), Stir Welded Joints on Airbus Aircraft Wing 9th International Conference on Fracture, pp. 77-79 Structure, 4th International Symposium Sidney, (1997), Vol. 1, pp. 523-530 25 Paris, P. C.; Sih, G. C.: Stress Analysis of on Friction Stir Welding, Park City, Utah 17 Prime, M. B.: Residual Stress Measurement Cracks, ASTM STP 381 (1964), pp. 30-81 (2003) by Successive Extension of a Slot: the 26 Tada, H.; Paris, P. C.; Irwin, G. R.: The 4 Lohwasser, D.: Application of Friction Stir Crack Compliance Method, Applied Mecha- Stress Intensity Analysis of Cracks Hand- Welding for Aircraft Industry, 2th Interna- nics Reviews 52 (1999), No. 2, pp. 75-96 book, 2th Edition, Paris Productions, (1985) tional Symposium on Friction Stir Welding, 18 Chiarelli, M. , A. Lanciotti, M. Sacchi: Fa- 27 Gallagher, J. P.: A Generalized Develop- Gothenburg, Sweden, (2000) tigue Resistance of MAG Welded Elements. ment of Yield Zone Models. AFFDL-TM-74- 5 Skorupa, M.: Load Interaction Effects Du- International Journal of Fatigue 21 (1999), 28-FBR, Wright Patterson Air Force Labo- ring Fatigue Crack Growth Under Variable pp. 1099-1110 ratory (1974) Amplitude Loading-A Literature Review, 19 Harter, J. A.: Afgrow, User’s Guide and 28 de Koning, A. U.; ten Hoeve, H. J.; Hendrik- Part I: Empirical Trends, Fatigue and Frac- Technical Manual, AFRL-VA-WP-TR-2002- sen, T. K.: The Description of Crack Growth ture of Engineering Materials and Structu- XXXX (2002) on the Basis of the Strip-Yield Model for res 21 (1998), pp. 987-1006 20 Forman, R. G.; Shivakumar, V.; Mettu, S. R.; Computation of Crack Opening Loads, 6 Skorupa, M.: Load Interaction Effects Du- Newman, J. C.: Fatigue Crack Growth Com- The Crack Tip Stretch and Strain Rates, ring Fatigue Crack Growth Under Variable puter Program „Nasgro“ Version 3, JSC- NLR-TP-97511, (1997) Amplitude Loading-A Literature Review, 22267B (2000) 29 Elber, W.: The Significance of Fatigue Crack Part II: Qualitative Interpratation. Fatigue 21 Ghidini, T.; Alfaro, U.; Dalle Donne, C.: Closure, ASTM STP 486 (1971), pp. 230-242 and Fracture of Engineering Materials and Fatigue Life of Friction Stir Welded Joints 30 Beghini, M.; Bertini, L.; Vitale, E.: Fatigue Structures, 22 (1998), pp. 905-926 in Presence of Corrosion Damage: Experi- Crack Growth in Residual Stress Fields: Ex- 7 Masubuchi, K.: Residual Stresses and ments and Calculations, Materialprüfung perimental Results and Modelling, Fatigue Distortion in Welded Aluminum Structures (Material Testing) 47 (2005), No. 7-8, and Fracture of Engineering Materials and and their Effects on Service Performance, pp. 456-461 Structures 17 (1994), No. 12, pp. 1433-1444 Welding Research Council Bulletins 172 (1972), pp. 1-30. 8 Dalle Donne, C.; Biallas, G.; Ghidini, T.; Abstract Raimbeaux, G.: Effect of Weld Imperfec- tions and Residual Stresses on the Fatigue Crack Propagation in Friction Stir Welded Abschätzungen von Ermüdungsrisswachstum bei Reibrühr- Joints, 2th International Conference on schweißnähten unter Flugbedingungen. Da sich das Reibrühr- Friction Stir Welding, Gothenburg, Sweden schweißen als Schlüsseltechnologie für primäre Luftfahrtstrukturen (2000) 9 Dalle Donne, C.; Lima, E.; Wegener, J.; erwiesen hat, sind Versuche und Modelle für den Ermüdungsriss- Pizalla, A.; Buslaps, T.: Investigations on fortschritt erforderlich. Allerdings gibt es nach wie vor einen Infor- Residual Stresses in Frciction Stir Welds. 3th International Conference on Friction mationsmangel bei dem Ermüdungsrisswachstum in FSW-Bauteile Stir Welding, Kobe, Japan (2001) unter variierender Last, sowohl bei einfacher Amplitudenvariation, 10 Dalle Donne, C.; Wegener, J.; Pyzalla, A.; Buslaps, T.: Investigations on Residual als auch unter Flugbedingungen. Die experimentellen Untersuchun- Stresses in Friction Stir Welds, 3th Interna- gen wurden an 4 mm dicken AA2024-T3 Basismetall- und FSW-Pro- tional Conference on Friction Stir Welding, ben durchgeführt, welche mit einem zentralen Riss versehen waren. Kobe, Japan (2001) 11 Ghidini, T.; Vugrin, T.; Dalle Donne, C.: Bei den FSW Proben wurde der Riss 5 mm außerhalb der Schweiß- Residual Stresses, Defects and Non-De- nahtmitte, im kritischsten Bereich der Fügung, plaziert. Alle Cou- structive Evaluation of FSW Joints, Welding International 19 (2004), No. 10, pp. 783-790 pons sind unter Lastsequenzen mit variierender Amplitude und 12 FALSTAFF, Descripiton of a Fighter unter einer standardisierten Flugsimulationssequenz Falstaff ge- Loading Standard for Fatigue Evaluation. prüft worden. Der Ermüdungsrissfortschritt wurde anschließend Combined Report LBF, NLR, IABG and F&W (1976) mit Hilfe der weit verbreiteten Luftfahrt-Bruchmechanik Software- 13 Zhang, S.; Marissen R.; Schulte, K.; Traut- pakete vorausgesagt. Variable Belastungen und schweißnahtbeding- mann, K. H.; Nowack, H.; Schijve, J.: Crack Propagation Studies on Al 7475 on the te Eigenspannungen wurden identifiziert und zuerst getrennt von Basis of Constant Amplitude and Selective einander simuliert. Die einzelnen Lösungen wurden dann über- Variable Amplitude Loading Histories, lagert, um das Ermüdungsrisswachstum von größeren FSW-Proben atigue and Fracture of Engineering Materi- als and Structures 10 (1987), No. 4, unter komplexen Lastspektren und Flugsimulation abzuschätzen. pp. 312-332

6 Jahrg. 48 (2006) 7-8 177 BETRIEBSFESTIGKEIT MP

The authors Claudia Polese was born in Cecina, Italy, Claudio Dalle Donne was born in Karlsruhe, in 1974. She received her M. Sc. Degree in Germany, in 1967. He received his Ph. D. in Tommaso Ghidini was born in Fidenza, Italy, Aeronautical Engineering from the Pisa Uni- Engineering Mechanics from the Karlsruhe in 1974. He received his M. Sc. Degree in versity in 2001. She’s now working in research University in 1996. He worked at the German Engineering Mechanics from the Parma Uni- activities regarding integral structures for Aerospace Centre of Cologne, at the Institute versity in 2000. He’s working at the German aeropace applications at the Department of of Materials Research as responsible for the Aerospace Centre of Cologne, at the Institute Aerospace Engineering of the University research projects on fatigue and fracture of Materials Research. His work is focused on of Pisa. mechanics as well as on friction stir welding. fatigue, fracture mechanics and numerical Agostino Lanciotti was born in Spoleto Since 2004 Dr. Dalle Donne is Senior Manager simulations of welded and not-welded structu- in 1949. He’s full Professor of Structural Metallic Structures at the EADS Corporate res. From June 2005 he works with Airbus Aerospace Technology at the Department of Research Center in Germany. Germany at the Fatigue and Damage Tolerance Aerospace Engineering of the University Department. of Pisa.

Jahrg. 48 (2006) 7-8 7 SELECTED PAPERS I

4.2.2 “FRICTION STIR WELDING” FOR AEROSPACE APPLICATIONS

[16] A. Lanciotti, C. Polese “Alcune proprietà di giunzioni “Friction Stir Welding” in due leghe di alluminio di impiego aerospaziale” Memoria presentata al Convegno “Friction Stir Welding per Applicazioni Aerospaziali”, Pisa, 1 Aprile 2003...... 179

[18] A. Lanciotti, C. Polese “Resistenza a fatica e caratteristiche di Damage Tolerance di giunzioni Friction Stir Welding nella lega alluminio-litio 2195-T8” Giornata di Studio: “Friction Stir Welding Day”, Torino, 15 Giugno 2004...... 193

[21] A. Lanciotti, R. Laudato, C. Polese “Proprietà meccaniche di giunzioni “Friction Stir Welding” in una lega alluminio-litio per applicazioni spaziali”. XVIII Conferenza della Associazione Italiana di Aeronautica e Astronautica, Volterra, 19-22 Settembre 2005. ..213

178 179

FRICTION STIR WELDING PER APPLICAZIONI AEROSPAZIALI, PISA 1 Aprile 2003

ALCUNE PROPRIETA' DI GIUNZIONI "FRICTION STIR WELDING" IN DUE LEGHE DI ALLUMINIO DI IMPIEGO AEROSPAZIALE

A. Lanciotti, C. Polese Dipartimento di Ingegneria Aerospaziale, Università di Pisa, Via G. Caruso, 56122 Pisa

Sommario

Nel presente lavoro si descrivono i risultati di una serie di prove mirate alla caratterizzazione di alcune proprietà di giunzioni realizzate con la tecnologia "Friction Stir Welding" in due leghe di alluminio di interesse aerospaziale; le leghe esaminate sono la tradizionale 2219-T851, una lega ampiamente utilizzata nel settore, e una lega alluminio-litio di nuova generazione. Gli aspetti investigati hanno riguardato la resistenza statica, la resistenza a tenso-corrosione, la resistenza a fatica, la propagazione dei difetti e le tensioni residue introdotte dal processo di saldatura. Il programma è stato completato con alcune prove comparative sui metalli base e su giunzioni realizzate con altre tecniche di saldatura, quali la TIG e la VPPA.

INTRODUZIONE

Le leghe alluminio-litio sono state studiate per oltre 50 anni ma solo recentemente hanno trovato importanti applicazioni nel settore aerospaziale. L'aggiunta di litio alle leghe di alluminio notoriamente aumenta la resistenza statica, aumenta il modulo elastico e diminuisce il peso specifico; queste proprietà sono basilari per il settore aerospaziale, che conseguentemente ha sempre mostrato interesse per tali materiali. Di fatto l'applicazione delle leghe alluminio-litio è stata limitata a causa di una o più proprietà negative, quali la bassa resistenza a fatica e la bassa resistenza a tenso- corrosione, particolarmente nella direzione trasversale corta (Short Transverse, ST). Questi problemi sono stati, almeno in gran parte, risolti nelle leghe di ultima generazione e quindi tali leghe, essendo anche saldabili e utilizzabili a temperature criogeniche, sono state utilizzate per la costruzione di serbatoi per combustibile di veicoli spaziali, in sostituzione di leghe più tradizionali, come la 2219, con risparmi ponderali considerevoli, [1]. In queste applicazioni solo la saldatura riesce a soddisfare i requisiti di ermeticità delle strutture e, fra le tecniche di saldatura, particolarmente interessante appare la "Friction Stir Welding, (FSW)", [2], una tecnica relativamente recente con la quale si realizzano giunzioni saldate senza raggiungere la temperatura di fusione del materiale. La giunzione viene realizzata attraverso il calore sviluppato per attrito da un perno che, ruotando, avanza lungo l’interfaccia fra due lamiere accostate di testa, come schematicamente rappresentato in fig. 1. Il minor apporto termico che compete a questa tecnica migliora notevolmente le proprietà meccaniche dei giunti saldati, mentre la minor temperatura di saldatura diminuisce la possibilità di contaminazione del metallo. Infine, praticamente scompaiono alcuni difetti tipici delle saldature quali porosità, cricche da ritiro e segregazioni. Allo stato attuale la saldatura delle leghe di alluminio è una realtà industriale, [3-7]; gli sviluppi futuri dovrebbero riguardare altri materiali strutturali di interesse, quali acciai e leghe di titanio e la realizzazione di cordoni di saldatura con sviluppo tridimensionale [8].

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Presso il Dipartimento di Ingegneria Aerospaziale dell'Università di Pisa, nell'ambito di una collaborazione con Alenia Spazio, è stato svolto un lavoro di caratterizzazione di giunzioni FSW in una lega alluminio-litio di nuova generazione e nella più tradizionale lega 2219-T851. L'attività sperimentale ha riguardato la resistenza statica, la resistenza a tenso-corrosione, la resistenza a fatica, la propagazione dei difetti e le tensioni residue introdotte nelle lastre dal processo di saldatura. Il programma è stato completato con alcune prove comparative effettuate sui metalli base e su giunzioni ottenute con altre tecniche di saldatura, quali la TIG per la lega alluminio-litio e la VPPA (Variable Polarity Plasma Arc) nella lega 2219-T851.

ATTIVITA' SPERIMENTALE

Materiali

Il materiale di partenza era costituito da piastre di 240x400 mm, spessore 5 mm, con saldatura longitudinale e da piastre di metallo base di 200x500 mm, spessore 5 mm. Le piastre di Al-Litio sono state ottenute per fresatura partendo da piastre di spessore maggiore, t=13 mm; per meglio interpretare alcuni risultati ottenuti nella sperimentazione, questo materiale è stato utilizzato per costruite alcune provette di metallo base mantenendo inalterato lo spessore. Tutti i materiali sono stati forniti da Alenia-Spazio, che ha anche effettuato le saldature. Tutte le prove sugli elementi saldati sono state effettuate nella condizione "as welded". La fig. 2 mostra una macro-fotografia di un cordone FSW nella lega Al-Litio; sono chiaramente identificabili la zona termo-meccanicamente alterata, con al centro il "nugget", circondata dalla zona termicamente alterata. Al centro del nugget i grani sono molto piccoli, mentre nel metallo base si osserva una struttura "stratificata", responsabile di una serie di problematiche che saranno descritte in seguito.

Attrezzature utilizzate e modalità di esecuzione delle prove.

Resistenza statica. La resistenza statica è stata misurata utilizzando provette di metallo base e provette FSW con saldatura trasversale. Le prove sono state effettuate con una macchina di trazione da 25 KN, misurando l’allungamento tramite un estensometro.

Tenso-corrosione. La resistenza a tenso-corrosione è stata valutata sulla base dei risultati di prove statiche comparative effettuate su gruppi di tre provini, di cui il primo gruppo (di confronto) sottoposto solo a prova statica, il secondo gruppo esposto ad ambiente corrosivo ed il terzo gruppo esposto allo stesso ambiente corrosivo e contemporaneamente tensionato. La prova, descritta in [9], è particolarmente rivolta alle leghe di alluminio, sia saldate che non saldate. L’ambiente corrosivo consiste nella immersione alternata, 10’ bagnato/50’ asciutto, dei provini in una soluzione al 3.5% di cloruro di sodio per un periodo di 30 giorni. Ai provini tenso-corrosi è stato applicato un carico tale da produrre una tensione pari al 75% della tensione di snervamento, misurata nei rispettivi provini di riferimento (non corrosi).

Fatica. Le prove di fatica sono state effettuate con un martinetto idraulico servocontrollato avente capacità di carico di 25 KN. In tutte le prove il rapporto di sollecitazione, R=Smin/Smax, era R=0.1.

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Propagazione di difetti. Sono state effettuate solamente prove di propagazione con difetto parallelo al cordone di saldatura. Per il difetto iniziale si sono investigate tre possibili locazioni: difetto al centro del cordone (bead) e difetto nella zona termicamente alterata, sia sul lato Advancing che sul lato Retreating (rispettivamente lato sul quale la velocità di rotazione e la velocità di traslazione dell'utensile si sommano oppure si sottraggono, vedi fig. 1). Durante le prove la dimensione dei difetti è stata misurata con un microscopio ottico 20x, montato su una slitta mobile dotata di un sistema di misura centesimale.

Tensioni residue. Per la misura delle tensioni residue agenti nelle lastre saldate si è fatto uso di una tecnica distruttiva che consiste nell'incollare alcuni estensimetri su una piastra che successivamente viene sezionata in modo da rilassare le tensioni interne. I tagli sono stati effettuati con una piccola fresa, avendo cura di proteggere gli estensimetri con sigillante in modo da evitare il contatto con piccoli trucioli metallici. Avanzamento e velocità di rotazione della fresa sono stati scelti in modo da minimizzare il riscaldamento della lastra, essendo sconsigliabile per questa misura l’uso di liquido refrigerante. Maggiori dettagli sono forniti in [10].

RISULTATI

Resistenza statica. Le prove statiche hanno riguardato la lega Al-Litio, sia metallo base che giunti FSW, mentre alcuni risultati erano già disponibili per la 2219-T851. I risultati ottenuti sono riportati nella fig. 3, unitamente ai risultati delle prove di tenso- corrosione, che saranno illustrati in seguito. Ciascun risultato presentato nella figura è la media dei risultati di tre prove nominalmente identiche. Per il metallo base la tensione di snervamento, Sy, e la tensione di rottura, Su, sono molto elevate, rispettivamente 549 e 594 MPa. Nei giunti FSW questi valori scendono a 277 e 418 MPa (50% e 70% dei corrispondenti valori del metallo base). Per il 2219-T851 prove analoghe [10] avevano fornito valori di tensione di snervamento e tensione di rottura di 338 e 453 MPa nel metallo base, e 165 e 265 MPa nei giunti saldati VPPA (48% e 58% dei corrispondenti valori del metallo base). Per i giunti in 2219-T851 FSW si dispone solo del valore della tensione di rottura [11] che risulta essere di 325 MPa (78% del valore del metallo base). Dal punto di vista della resistenza meccanica la lega Al-Litio ha quindi proprietà decisamente più elevate rispetto alla tradizionale lega 2219-T851 e nella lega 2219 i giunti FSW hanno resistenza più elevata dei giunti VPPA. Sempre relativamente alla lega Al-Litio, si è misurato un allungamento a rottura del 5.5% nel metallo base, mentre nel metallo saldato questa proprietà non è immediatamente definibile, se non premettendo alcune considerazioni: la rottura dei giunti saldati avviene ad una tensione più bassa della tensione di snervamento del metallo base per cui, quando il giunto saldato si rompe, il materiale fuori dal cordone di saldatura è ancora in campo elastico. Conseguentemente per ottenere un valore significativo, l'allungamento deve essere misurato solo all'interno del cordone di saldatura. A tal fine su un provino saldato sono state tracciate delle linee di riferimento in senso trasversale, distanziate di circa 2 mm. L'esatta distanza fra le linee è stata misurata con un microscopio ottico 100x, equipaggiato con una slitta mobile con lettura millesimale dello spostamento. Il provino è stato quindi sollecitato con una carico tale da produrre una tensione pari al 98% della tensione di rottura. Misurando nuovamente la distanza fra le linee si sono misurati allungamenti permanenti fino al 15%, molto superiori quindi rispetto al valore, 5.5%, rilevato nel metallo base. La fig. 4 mostra il risultato delle misure di allungamento effettuate in otto punti lungo il cordone di

- 3 - 182 saldatura; localmente i valori sono elevati, ma, poichè tale allungamento è limitato al solo cordone di saldatura, si ha spesso nelle saldature l'errata impressione di rotture fragili.

Resistenza a tenso-corrosione. Le prove di tenso-corrosione hanno riguardato la lega Al-Litio, sia metallo base che giunti FSW. In entrambi i casi si è osservata una elevatissima resistenza a tenso-corrosione, come mostrato nella fig. 3, nella quale sono riportati i risultati delle prove statiche condotte su provette esposte ad ambiente corrosivo per 30 giorni e provette esposte allo stesso ambiente corrosivo, contemporaneamente sollecitate con un carico tale da produrre una tensione pari al 75% della tensione di snervamento misurata sulle provette di riferimento (non corrose). L'esposizione all'ambiente corrosivo, anche in presenza di sollecitazione, non riduce la resistenza statica; addirittura per il metallo base si osserva un leggero aumento della tensione di snervamento. Questo fenomeno, "environmetal hardening", osservato anche da altri ricercatori, [12], può essere associato ad una diminuzione della duttilità e della Fracture Toughness. Nel caso in esame il contemporaneo aumento dell'allungamento a rottura sembra escludere questa possibilità. Dopo l'esposizione ad ambiente corrosivo non si sono osservati pits di corrosione nella lega Al-Litio, per cui si può concludere che in questo materiale anche la resistenza a corrosione generalizzata è molto elevata; prove effettuate in precedenza sul 2219-T851 metallo base e giunzioni VPPA, [10], mostravano una elevata resistenza a tensocorrosione ma la presenza di pits, particolarmente sulle provette tenso-corrose, era molto comune.

Resistenza a fatica. Sono state effettuate prove di fatica sulla lega Al-Litio su provette costruite con metallo base, provette FWS e provette saldate con processo TIG (Tungsten Inert Gas); inoltre erano disponibili i risultati di prove effettuate precedentemente su provette in 2219-T851 saldate VPPA. Tutte le prove sono state effettuate in aria, carico sinusoidale, rapporto di sollecitazione R=Smin/Smax=0.1. I risultati ottenuti sono mostrati nella fig. 5. Particolarmente nel campo degli alti numeri di cicli, la resistenza a fatica della lega Al-Litio è molto elevata, come si può dedurre confrontando i risultati ottenuti con quelli della lega 7075-T6 [13], che, fra le leghe più convenzionali, è quella che presenta la maggior resistenza. La resistenza a fatica è molto elevata anche per le giunzioni FSW, mentre tecniche di saldatura convenzionali, come la saldatura TIG, non appaiono adeguate dal punto di vista della resistenza a fatica di questo materiale. Per quanto riguarda la lega 2219-T851, la saldatura VPPA è considerata una buona tecnica di saldatura ed è stata utilizzata per la costruzione di importanti strutture spaziali; i risultati delle prove effettuate hanno comunque mostrato una resistenza a fatica più bassa di quella delle giunzioni FSW nella lega Al-Litio. Le rotture per fatica nei giunti FSW si sono tipicamente manifestate nella zona terminale del contatto fra lo spallamento dell'utensile e la piastra, fig. 6. Per quanto riguarda le provette in Al-Litio metallo base, si sono osservate delle rotture atipiche, caratterizzate da uno sfaldamento in senso longitudinale delle provette, come mostrato nella fig. 7. In alcune provette si è monitorizzata l'evoluzione di queste cricche e si è potuto osservare che queste si manifestano solo nelle ultimissime fasi della vita a fatica; la fig. 8 mostra alcune fotografie scattate durante una di queste prove. I dati della prova erano: Smax=300 MPa; R=0.1; numero di cicli a rottura 125200. Dopo circa 110000 cicli si è osservata la prima cricca per fatica che a 121000 cicli aveva raggiunto una dimensione di circa 2.5 mm, fig. 8a. A questo punto la cricca si è ramificata in senso longitudinale, come mostrato nella fig. 8b. La propagazione in direzione longitudinale è proseguita unitamente alla propagazione in direzione trasversale, fig. 8c, fino alla rottura finale. Il comportamento osservato è stato messo in relazione con la struttura stratificata che il materiale presentava, come già illustrato

- 4 - 183 nella fig. 2 (dettaglio sul lato destro della figura). Per meglio approfondire questo aspetto sono state costruite alcune provette orientate nella direzione trasversale corta del materiale, fig. 9, utilizzando le piastre aventi spessore di 13 mm. Il provino ovviamente non è di dimensioni standard, ma lo spessore delle piastre non consentiva un disegno più soddisfacente, volendo mantenere una sezione resistente non piccolissima per problemi legati alle attrezzature di prova. Nonostante le ridotte dimensioni, tale provino si è dimostrato efficace, anche se alcune rotture per fatica si sono manifestate nelle filettature di afferraggio; il problema è stato successivamente risolto riducendo il diametro esterno da 6 a 5.5 mm. Il foro centrale, necessario per poter tornire le provette a sbalzo data la loro ridotta lunghezza, è risultato anche essere un utile stratagemma per ridurre la concentrazione di tensione nella sezione di prova. Infatti mentre per la provetta piena si è valutato con un modello agli elementi finiti una concentrazione di tensione di 1.29, per la provetta con foro centrale tale valore si riduce a 1.13. I risultati delle prove effettuate sono riportati nella fig. 10. Come ipotizzato, le prove hanno mostrato una minor resistenza a fatica nella direzione trasversale corta rispetto alla direzione longitudinale e questa differenza aumenta con il numeri di cicli. Lo scatter nei risultati è elevato, ma paragonabile con quello ottenuto per le provette costruite nella direzione longitudinale del materiale. Sulle superfici di frattura si osserva un continuo cambiamento del piano di giacitura delle rotture per fatica, fig. 11, che sembrano incunearsi nella stratificazione del materiale, già osservata in fig. 2. Globalmente i risultati delle prove di fatica per la lega Al-Litio mostrano un materiale molto resistente ma che dovrebbe essere ulteriormente migliorato, particolarmente per quanto concerne la resistenza nella direzione ST. Non è stato possibile, dato il piccolo spessore delle piastre, costruire provette per la misura della Fracture Toughness nella direzione trasversale corta; anche questo aspetto meriterebbe un approfondimento.

Propagazione di difetti. I risultati delle prove di propagazione di difetti sono riportati nella fig. 12 per la lega Al-Litio e in fig. 14 per la 2219-T851 nella forma velocità di propagazione di un difetto, da/dN, in funzione della variazione nel ciclo di carico del fattore di intensità degli sforzi, DK. Relativamente alla lega Al-Litio si osserva per il metallo base un ginocchio nella curva di propagazione nella zona di valori DK=5¸8 MPa m . Tale anomalia scompare nelle curve di propagazione del metallo saldato. Per i pannelli saldati il difetto iniziale è stato introdotto in tre possibili locazioni: al centro del cordone (bead) e nella zona termicamente alterata, sia sul lato Advancing che sul lato Retreating. In tutti i casi i difetti si sono propagati mantenendo la stessa giacitura iniziale, per cui non sembra esistere nel cordone di saldatura una locazione critica. Le due locazioni nella zona termicamente alterata sono equivalenti per quanto concerne la velocità di propagazione di un difetto e tale velocità è uguale a quella misurata nel metallo base; al centro del cordone si osserva, viceversa, un leggero aumento della velocità di propagazione. Anche nelle prove di propagazione della lega Al-Litio, metallo base, si è evidenziata la struttura stratificata del materiale, che in questo caso si è manifestata come un aspetto "legnoso" delle superfici di frattura per le piastre aventi orientamento TL, fig. 13. Confrontando i risultati ottenuti con quelli già disponibili per il 2219-T851, [10], si osserva, per i metalli base, un minore velocità di propagazione per la lega Al-Litio, esclusi i valori più bassi di DK, fig. 12. Per le giunzioni FSW nella lega 2219-T851 si è osservata una maggior velocità di propagazione per tutti i valori di DK rispetto al metallo base; i risultati ottenuti non dipendono dalla locazione del difetto iniziale nel cordone di saldatura. Risultati disponibili per giunzioni in 2219-T851 saldate VPPA [10] mostravano una velocità di propagazione più bassa, mediamente comparabile con quella del metallo base.

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Tensioni residue. Le fig. 15 e 16 mostrano per i due materiali, lega Al-Litio e 2219- T851, l'andamento della tensione residua agente nella direzione del cordone di saldatura, in funzione dell'ascissa trasversale, avente origine nel centro del cordone stesso, come mostrato nell'inserto in fig. 15. Per la lega Al-Litio si raggiunge una tensione residua massima di circa 140 MPa nella zona terminale del contatto fra utensile e pezzo. La distribuzione di tensioni è asimmetrica, con un picco di tensione leggermente più elevato dal lato Retreating. Per il 2219-T851 la distribuzione di tensioni residue mantiene la forma a W rovesciato, tipica delle leghe di alluminio, ma si raggiungono valori massimi decisamente più bassi, circa 85 MPa. Per questo materiale si disponeva di risultati simili relativi a saldature VPPA in lastre aventi uno spessore di 7 mm, [10], e riportati nella stessa figura. In questo caso il valore massimo della tensione residua era di circa 140 MPa e la zona interessata da importanti tensioni residue di trazione era più ampia che nel caso di FSW. Si può quindi concludere che la minor temperatura alla quale avviene il processo FSW, rispetto alle saldature per fusione, ha un effetto benefico anche sull'entità e sulla distribuzione delle tensioni residue. E' interessante rilevare che, nelle due giunzioni FSW, il rapporto fra il valore massimo della tensione residua e la tensione di snervamento del materiale base è praticamente uguale, 25.5% nella lega Al-Litio e 25.1% nel 2219-T851. Da notare infine che la distribuzione di tensioni residue mostrata in fig. 16 per le saldature VPPA è evidentemente non equilibrata, ma non si è trovata una giustificazione plausibile per quanto osservato.

CONCLUSIONI

La “Friction Stir Welding” è una tecnica di saldatura relativamente recente che ha esteso alle lamiere i vantaggi della saldatura per attrito, una saldatura che avviene allo stato solido, con minor esposizione del metallo a possibili contaminazioni, rispetto alle saldature convenzionali; inoltre praticamente scompaiono alcuni difetti tipici delle saldature quali porosità, cricche da ritiro e segregazioni. Nell'ambito di una collaborazione con Alenia Spazio, è stata effettuata una caratterizzazione di alcune proprietà meccaniche di due leghe di alluminio di interesse aerospaziale: la tradizionale 2219-T851, una lega ampiamente utilizzata nel settore e una lega alluminio-litio di nuova generazione. Gli aspetti investigati hanno riguardato la resistenza statica, la resistenza a tenso-corrosione, la resistenza a fatica, la propagazione dei difetti e le tensioni residue introdotte dal processo di saldatura. Il programma è stato completato con alcune prove comparative sui metalli base e su giunzioni realizzate negli stessi materiali con altre tecniche di saldatura, quali la TIG e la VPPA. Il programma di prove svolto ha consentito di trarre le seguenti conclusioni:

La lega Al-Litio ha resistenza statica molto elevata e mantiene questa proprietà anche nelle giunzioni FSW. La resistenza delle giunzioni FSW nel 2219-T851 è maggiore di quella delle giunzioni VPPA. Per la lega Al-Litio l'allungamento a rottura è del 5.5% nel metallo base mentre nei giunti FSW si sono misurati, limitatamente al cordone di saldatura, allungamenti anche del 15%. La rottura dei giunti saldati avviene ad una tensione inferiore alla tensione di snervamento del metallo base per cui la deformazione permanente risulta essere limitata al solo cordone di saldatura.

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La lega Al-Litio, sia metallo base che giunti FSW, presenta elevata resistenza a tenso- corrosione e a corrosione generalizzata.

La resistenza a fatica della lega Al-Litio è elevata sia per il metallo base che per le giunzioni FSW. Decisamente inferiore è stato il risultato delle giunzioni TIG; evidentemente il maggior apporto di calore necessario per realizzare la saldatura è molto dannoso in questi materiali. La lega Al-litio presentava una struttura stratificata nella direzione trasversale corta; questa struttura, osservabile al microscopio, ha prodotto rotture per fatica anomale nelle provette di metallo base, con formazione di cricche secondarie che si propagavano nella direzione del carico. Prove di fatica effettuate su provette ricavate nella direzione trasversale corta hanno evidenziato una minor resistenza in tale direzione. Per il 2219-T851 la resistenza a fatica di giunzioni VPPA è inferiore rispetto a quella delle giunzioni FSW.

La velocità di propagazione di un difetto nelle piastre in Al-Litio FSW è leggermente più elevata al centro del cordone, mentre nella zona termicamente alterata, sia nel lato Advancing che nel lato Retreating, la velocità di propagazione è più bassa e uguale a quella misurata nel metallo base. Relativamente ai metalli base, la velocità di propagazione è minore nella lega Al-Litio rispetto alla 2219-T851, esclusi i valori più bassi di DK. Anche nelle prove di propagazione nelle piastre di metallo base Al-Litio, si è evidenziata la struttura stratificata del materiale, che in questo caso si è manifestata come un aspetto "legnoso" delle superfici di frattura per le piastre aventi orientamento TL.

Per la lega 2219-T851, giunti FSW, si osserva una maggior velocità di propagazione per tutti i valori di DK rispetto al metallo base; la velocità di propagazione nel metallo saldato non dipende dalla locazione del difetto iniziale nel cordone di saldatura. Risultati precedenti mostravano per la lega 2219-T851 VPPA una velocità di propagazione più bassa, mediamente comparabile con quella del metallo base.

Nella lega Al-Litio il processo FSW produce una tensione residua massima di circa 140 MPa. Per il 2219-T851 si raggiunge un valore massimo decisamente più basso, pari a circa 85 MPa. Confrontando i risultati ottenuti per il 2219-T851 con alcuni risultati disponibili relativi a saldature VPPA nello stesso materiale, si osserva una tensione residua massima più elevata, circa 140 MPa nel caso VPPA ed inoltre la zona interessata da importanti tensioni residue di trazione è più ampia. Si può quindi concludere che la minor temperatura alla quale avviene il processo FSW, rispetto alle saldature per fusione, ha un effetto benefico anche sull'entità e sulla distribuzione delle tensioni residue.

BIBLIOGRAFIA

[1] M. Braukus, J. Malone "Shuttle Super Lightweight Fuel Tank Completes Test Series". NASA Press realise July 1996. [2] W.M.Thomas, E.D. Nicholas, J.C. Needham, M.G. Church, P. Templesmith, C.J. Dawes (1991). International Patent Application No. PCT/IGB92/02203 and GB Patent Application No. 9125978.9. [3] R. Talwar, D. Bolser, R. Lederich, J. Baumann "Friction stir welding of airframe structures". Proceedings of the 2nd International Symposium on Friction Stir Welding, Gothenburg, Sweden, 2001.

- 7 - 186

[4] D. Waldron "Application of Friction Stir Welding for Delta Rocket Fuel Tanks". Proceedings of the 1st International Symposium on Friction Stir Welding, Thousand Oaks, California, USA, 1999. [5] I. Henderson "Exploiting friction stir welding in explosively formed aluminium boat hull construction". Proc. Conf. 'INALCO' 98. 7th International Conference on Joints in Aluminium. Cambridge, UK, 1998, pp. 251-257. [6] T. Kawasaki, T. Makino, S. Todori, H. Takai, M. Ezumi, Y. Ina "Application of friction stir welding to the manufacturing of next generation "A-train" type rolling stock". Proc. 2nd International Symposium on Friction Stir Welding, Gothenburg, Sweden, 2000. [7] M.R. Johnsen "Friction stir welding takes off at Boeing". Welding Journal 78 (2), 1999, pp. 35-39. [8] A. Strombeck, C. Schilling, J.F. dos Santos "Robotic friction stir welding -Tool, Technology and Applications". 2nd International Symposium on Friction Stir Welding, Gothenburg, Sweden, 2000. [9] Anon. "Determination of the susceptibility of metals to stress corrosion cracking". European Space Agency, Spec. n. PSS-01-737, Issue 1, 1981. [10] A. Lanciotti, A. Belmondo "Mechanical properties of Al 2219 VPPA weldments". 48th International Astronautical Congress, Turin, 1997. Paper IAF-97-1.4.07. [11] R. Laudato, A. Belmondo, F. Giuggioli, M. Armaiuoli "Applicazione della tecnologia Friction Stir Welding alla saldatura di leghe di alluminio di impiego aerospaziale" Giornate Italiane della Saldatura, Milano 7-9 Novembre 2001. [12] W. Hu, E.I Meletis "Corrosion and environment-assisted cracking behaviour of friction stir welded Al 2195 and Al 2219 alloys". Material Science Forum, Vol.331- 337, pp. 1683-1688, 2000. [13] Anon "Metallic Materials and Elements for Aerospace Vehicle Structures". Military Handbook 5E. June 1986.

Ringraziamenti

Gli autori ringraziano il Sig. Guido Becchi e l'Ing. Nicola Morgantini per l'aiuto fornito nella esecuzione delle prove sperimentali.

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SPALLAMENTO PERNO

ADVANCING SIDE

RETREATING SIDE Fig. 1 - Rappresentazione schematica del processo Friction Stir Welding.

2 mm

25 mm

Fig. 2 - Macrografia di un cordone FSW nella lega Al-Litio

25 mm

800 A% Metallo base 700 7 Riferimento Corrosione 600 6 Tenso-corrosione FSW 500 5

400 4

300 3 S (MPa) 200 2

100 1

0 0 2S 4y 6 8Su 10 12A% 14 16Sy 18 20Su 22

Fig. 3 - Risultati delle prove di tenso-corrosione effettuate sulla lega Al-Litio. - 9 - 0.20 188 S=412 MPa

0.15

0.10 ) l / 0.05 l D ( p e 0.00

-0.05 Cordone di saldatura

-0.10 -20 -15 -10 -5 0 5 10 15 20 x (mm)

Fig. 4 - Deformazione plastica nel cordone di saldatura della lega Al-Litio misurata durante una prova di trazione interrotta.

400 7075-T6 [13]

Al-Litio Metallo base 300 Smax (MPa)

Al-Litio FSW 200 2219-T851 VPPA

Al-Litio TIG

R=0.1 100 104 105 106 107 Cicli

Fig. 5 - Risultati delle prove di fatica.

Fig. 6 - Rottura per fatica nella zona termicamente alterata. Lega Al-Litio. - 10 - 189

Fig. 7 - Tipiche rotture per fatica nella lega Al-Litio.

Smax=300 MPa; Nf=125200 t = 5 mm

a) b) c) N=121000 N=123800 N=124700

Fig. 8 - Evoluzione di una rottura per fatica nella lega Al-Litio.

6 L 4 LT

R2 ST 13

M9 Fig. 9 - Provino per prove statiche e per prove di fatica nella direzione trasversale corta della lega Al-Litio

- 11 - 190

400

300

200

R=0.1 Smax (MPa)

Al-Litio L Al-Litio ST

100 104 105 106 107 Cicli Fig. 10 - Lega Al-Litio. Risultati delle prove di fatica; provette in direzione L e ST.

0.5 mm

Fig. 11 - Lega Al-Litio. Superficie di frattura nella direzione ST.

- 12 - 191

0.01

1E-3

1E-4

1E-5

1E-6

da/dN [mm/ciclo] Al-Litio; Metallo base Al-Litio; FSW; [Advancing; Retreating] 1E-7 Al-Litio; FSW; Bead 2219-T851; Metallo base 1E-8 3 4 5 6 7 8 9 10 20 30 40 50 60 70 DK [MPa m]

Fig. 12 - Risultati delle prove di propagazione. Lega Al-Litio; metallo base e saldature FSW.

TL

LT

Fig. 13 - Lega Al-Litio; superfici di frattura. Direzioni LT e TL.

10-2

10-3

10-4 da/dN [mm/ciclo]

10-5

10-6 2219-T851 Base Material 2219 Variable Polarity Plasma Arc Welding 10-7 2219 FSW [HAZ, Advancing] 2219 FSW [HAZ, Retreating] 2219 Fsw [Bead]

10-8 1 2 3 4 5 6 7 8 910 20 30 40 50 60 DK [MPa m] Fig. 14 - Risultati delle prove di propagazione. Lega 2219-T8; metallo base e saldature FSW e VPPA. - 13 - 192

160

140 Lega AL-Litio

120 100 y 80

60 x

40

Sres [MPa] 20

0

-20 Cordone di saldatura -40

-80 -60 -40 -20 0 20 40 60 80 x [mm]

Fig. 15 - Tensioni residue agenti nella direzione longitudinale. Materiale: Lega Al-Litio, processo FSW.

140 Mat.: 2219-T851 120

100 Cordone VPPA FSW 80 VPPA

60

40

20 Sres [MPa]

0 Cordone FSW

-20

-40

-100 -80 -60 -40 -20 0 20 40 60 80 100 x [mm]

Fig. 16 - Tensioni residue agenti nella direzione longitudinale. Materiale: 2219-T851, processi FSW e VPPA.

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Alenia

SPAZIO

A FINMECCANICA COMPANY

RESISTENZA A FATICA E CARATTERISTICHE DI DAMAGE TOLERANCE DI GIUNZIONI FRICTION STIR WELDING NELLA LEGA ALLUMINIO-LITIO 2195-T8

A. LANCIOTTI, C. POLESE Dipartimento di Ingegneria Aerospaziale, Università di Pisa Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno194 2004 - 1 -

RESISTENZA A FATICA E CARATTERISTICHE DI DAMAGE TOLERANCE DI GIUNZIONI FRICTION STIR WELDING NELLA LEGA ALLUMINIO-LITIO 2195-T8

A.Lanciotti, C.Polese1

Dipartimento di Ingegneria Aerospaziale, Università di Pisa, Via G. Caruso, 56122 Pisa

Sommario

Si descrivono i risultati di una serie di prove effettuate per la caratterizzazione delle proprietà generali e delle proprietà di Damage Tolerance di giunzioni realizzate con la tecnologia "Friction Stir Welding" nella lega alluminio-litio 2195-T8. L’attività, tutt’ora in corso, viene svolta nell’ambito di una collaborazione fra il Dipartimento di Ingegneria Aerospaziale dell’Università di Pisa e la Alenia Spazio, ed è mirata alla qualifica del processo di saldatura. Gli aspetti investigati riguardano la resistenza statica, la resistenza a fatica, la propagazione dei difetti e le tensioni residue introdotte dal processo di saldatura. In corso di svolgimento sono le prove di resistenza a tenso- corrosione e di resistenza a frattura. I risultati delle prove, confrontati con gli stessi risultati ottenuti dal metallo base, hanno mostrato le ottime proprietà dei giunti saldati.

INTRODUZIONE

Le leghe alluminio-litio, per quanto siano state studiate per oltre 50 anni, solo recentemente hanno trovato importanti applicazioni nel settore aerospaziale. L'aggiunta di litio alle leghe di alluminio notoriamente aumenta la resistenza statica e il modulo elastico, diminuendo contemporaneamente il peso specifico; tutte queste proprietà sono basilari per il settore aerospaziale, che conseguentemente ha sempre mostrato interesse per questi materiali. Dal punto di vista pratico, l'applicazione delle prime leghe alluminio-litio, quali la 2020, la 2090, la 2091 e la 8090, è stata limitata a causa di una o più proprietà negative, quali la bassa resistenza a fatica, la bassa resistenza a tenso-corrosione e a frattura, particolarmente nella direzione trasversale corta (Short Transverse, ST), unitamente ad una spiccata anisotropia. Questi problemi sono stati, almeno in gran parte, risolti nelle leghe di ultima generazione, caratterizzate da un minor contenuto di litio; tali leghe, essendo anche saldabili e utilizzabili a temperature criogeniche, sono state utilizzate per la costruzione di serbatoi per combustibile di veicoli spaziali, in sostituzione di leghe più tradizionali, come la 2219. I risparmi ponderali sono stati considerevoli poiché le leghe alluminio-litio, rispetto alle leghe tradizionali, presentano una resistenza statica superiore del 50%, un modulo elastico superiore del 5% e un peso specifico inferiore del 5%, [1]. In queste applicazioni solo la saldatura riesce a soddisfare i requisiti di ermeticità delle strutture; fra le tecniche di saldatura, particolarmente interessante appare la "Friction Stir Welding, (FSW)", [2], una tecnica relativamente recente sviluppata dal TWI (The Welding Institute), con la quale si realizzano giunzioni saldate senza raggiungere la temperatura di fusione del materiale. La giunzione viene realizzata attraverso il calore sviluppato per attrito da un perno che, ruotando, avanza lungo l’interfaccia fra due lamiere accostate di testa. Il minor apporto termico che compete a questa tecnica migliora notevolmente le proprietà meccaniche dei giunti saldati, mentre la minor temperatura di saldatura diminuisce la possibilità di contaminazione del metallo. Infine, praticamente scompaiono alcuni difetti tipici delle saldature quali porosità, cricche da ritiro e segregazioni. Allo stato attuale la saldatura

1 [email protected] [email protected] Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno195 2004 - 2 - delle leghe di alluminio è una realtà industriale, [3-7]; gli sviluppi futuri dovrebbero riguardare altri materiali strutturali di interesse, quali acciai e leghe di titanio e la realizzazione di cordoni di saldatura con sviluppo tridimensionale, [8]. Le problematiche di fatica, frattura e corrosione di giunzioni realizzate con Friction Stir Welding in una lega alluminio-litio di ultima generazione, la 2195-T8, sono l’oggetto di una attività di ricerca congiunta, tutt’ora in corso di svolgimento, fra il Dipartimento di Ingegneria Aerospaziale dell’Università di Pisa e la Alenia Spazio. Scopo della collaborazione è la caratterizzazione, finalizzata alla qualifica del processo di saldatura, delle proprietà di giunzioni FSW realizzate dalla Alenia Spazio. Su tali problematiche era stata precedentemente svolta una attività di caratterizzazione di giunzioni realizzate durante la fase di sviluppo preliminare del processo di saldatura, [9]; le giunzioni oggetto della presente indagine si differenziano dalle precedenti, per il diverso sistema di controllo della posizione verticale dell’utensile durante la realizzazione del giunto, basata sulla forza mutua fra utensile e pezzo e non sullo spostamento verticale, oltre che per un aggiustamento dei parametri di saldatura e una diversa geometria dell’utensile, che consente la effettuazione di buone saldature senza richiedere l’inclinazione dell’utensile stesso nella direzione di avanzamento longitudinale. Operando in questo modo si limita la diminuzione di spessore che in alcuni casi si può osservare in giunti FSW. Per lo svolgimento di questa attività il Dipartimento ha usufruito di un finanziamento del Ministero dell’Istruzione, dell’Università e della Ricerca, programma COFIN 2003/2004. Il programma dell’attività riguarda diversi temi quali: Resistenza a fatica. Tensioni residue in giunti saldati. Propagazione di difetti. Suscettività a tenso-corrosione. Resistenza a frattura. Il lavoro fino ad oggi svolto ha riguardato tutti gli argomenti, esclusa la suscettività a tenso-corrosione e la resistenza a frattura. Sul secondo argomento non si dispone di risultati mentre l’attività precedentemente svolta, [9], aveva comunque mostrato una elevata resistenza a tenso corrosione delle giunzioni FSW nel 2195-T8.

ATTIVITA’ SPERIMENTALE

Giunti saldati

Il materiale reso disponibile dalla Alenia Spazio era costituito da piastre piane saldate aventi dimensioni 300*850 mm, spessore 5 mm, ottenute saldando due piastre di 150*850 con direzione delle fibre longitudinale. Una piastra di dimensioni maggiori, 560*900 mm, sempre con saldatura longitudinale e spessore 5 mm, è stata utilizzata per la misura delle tensioni interne. Inoltre erano disponibili alcune piastre piane e alcune piastre curve (raggio 440 mm) contenenti ciascuna tre tratti di saldatura con doppia passata intervallati con tre chiusure del cordone di saldatura. La doppia passata si rende necessaria nel caso di saldatura circonferenziale di recipienti cilindrici, nei quali la saldatura necessariamente inizia e termina nella struttura stessa. In questi casi si utilizza un particolare utensile, il cui perno può essere estratto dal pezzo saldato, mantenendo in contatto la spalla. Operando in questo modo la saldatura termina senza lasciare il foro tipico dell’estrazione dell’utensile al termine della saldatura Friction. Sia la saldatura in doppia passata che il punto di chiusura della saldatura sono stati analizzati dal punto di vista del comportamento a fatica. Tutta l’attività di caratterizzazione è stata condotta su piastre “as welded”, essendo questa la tipica condizione in cui sono utilizzate le strutture saldate di impiego spaziale.

Resistenza statica

La resistenza statica è stata misurata utilizzando provette di metallo base e provette FSW con saldatura trasversale. Le prove statiche sono state condotte impiegando una macchina Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno196 2004 - 3 - servoidraulica con capacità di carico di 25 KN, misurando l’allungamento dei provini durante le prove tramite un estensometro. I soli valori medi dei risultati ottenuti in 3 prove nominalmente identiche sono riportati in fig. 1. Per il metallo base si è ottenuto un valore medio della tensione di snervamento Sy pari a 549 MPa e una tensione di rottura Su pari a 594 MPa con un allungamento a rottura del 5.5%. Per i giunti FSW si è ottenuto per lo snervamento un valore di 277 MPa e la rottura finale a 418.25 MPa, pari rispettivamente al 50 e 70% dei corrispondenti valori del metallo base, [9]. Il parametro maggiormente influenzato dalla saldatura è la tensione di snervamento e su tale valore dovrebbe quindi essere valutato “il coefficiente del giunto”. Per quanto riguarda l’allungamento percentuale, come già descritto nel lavoro precedente, [9], si è misurato un allungamento medio nel tratto interessato dal cordone di saldatura del 10.5%. Poichè la tensione di rottura del cordone è inferiore alla tensione di snervamento del materiale base, si ha collasso plastico nel cordone di saldatura quando il materiale all’esterno del cordone è ancora in campo elastico. Si può quindi erroneamente ritenere fragile la rottura di un giunto saldato, quando in realtà questa è una rottura duttile, ma con plasticizzazione limitata al solo cordone di saldatura.

Microdurezza

Le misure di microdurezza, effettuate presso i laboratori di Alenia Spazio, mostrano una diminuzione della durezza in corrispondenza del cordone. Il valore di durezza Vickers (HV0.1) del metallo base è pari a 154, mentre al centro del cordone si osserva una riduzione media del 37%. L’andamento del profilo di durezza è molto simile sia nel caso di singola che doppia passata, con una maggiore uniformità dei risultati nel caso della doppia passata. La riduzione di durezza, e la conseguente diminuzione di proprietà meccaniche, è concentrata nella zona terminale della zona termomeccanicamente alterata, zona in cui si sono tipicamente manifestate sia le rotture statiche che di fatica, con una diminuzione leggermente più accentuata in corrispondenza del lato retreating. È stato dimostrato sperimentalmente, [10], che ad un aumento della durezza nella zona termicamente alterata, ottenibile entro certi limiti agendo sui parametri di saldatura, corrisponde un aumento di resistenza meccanica dei giunti.

Sezioni metallografiche dei giunti saldati

Alenia Spazio ha realizzato una serie di macrografie dei giunti saldati che mostrano chiaramente le classiche caratteristiche di una giunzione FSW: il nugget centrale, caratterizzato da una grana fortemente equiassica, la zona termomeccanicamente alterata (TMZ) e la zona termicamente alterata (HAZ). Sono altresì visibili i cosiddetti “onion rings” al centro del nugget e una elevata fibrosità del materiale base, già osservata nel precedente lavoro. L’aspetto del cordone è visibilmente diverso nei due casi di singola e doppia passata: con una seconda passata si ha una maggiore uniformità del giunto, aumentano le dimensioni del nugget e diventa quindi molto più brusco il passaggio tra zona termomeccanicamente alterata e metallo base. Inoltre da analisi micrografiche appare una grana del nugget centrale più fine con una maggiore dispersione degli ossidi rispetto al cordone a singola passata.

Resistenza a fatica

La resistenza a fatica non è in genere una proprietà di interesse per il progetto di strutture spaziali, per le quali domina un dimensionamento basato sui criteri della meccanica della frattura. L’effettuazione di prove di fatica è comunque interessante per poter stabilire il punto più debole dei cordoni di saldatura, nel quale inserire i difetti iniziali per le prove di meccanica della frattura. Nel lavoro precedentemente svolto sui giunti FSW realizzati nella fase di messa a punto del processo, si era osservata una elevata resistenza a fatica rispetto alle giunzioni saldate Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno197 2004 - 4 - con altre tecniche, come il TIG (Tungsten Inert Gas) e il VPPA (Variable Polarity Plasma Arc). Nell’attività in corso sono state effettuate prove di fatica di giunti FSW impiegando provini tipo dogbone con saldatura trasversale. Le prove hanno riguardato saldature realizzate con singola passata e con doppia passata; alcuni provini, sempre con saldatura trasversale, sono stati costruiti in modo da contenere al centro un punto di chiusura di un cordone di saldatura realizzato nel modo descritto precedentemente (sollevamento progressivo del perno mentre l’utensile ruota e avanza, mantenendo la spalla sempre in contatto con il pezzo in lavorazione). Tutte le prove sono state condotte con un martinetto idraulico servocontrollato avente una capacità di carico di 50 KN. In tutte le prove il rapporto di sollecitazione R=Smin/Smax era pari a 0.1. I risultati ottenuti sono riportati in fig. 4; come si può osservare i risultati relativi al metallo base presentano una elevata dispersione e questa caratteristica si mantiene nei giunti saldati con passata singola; per questo gruppo di risultati una linea di best fit costituita da due segmenti sembra interpretare meglio il fenomeno. Una elevata dispersione era stata osservata anche nei risultati della prima serie di prove, condotte su giunti saldati con processo non qualificato. Purtroppo, anche se consigliata come buona norma nel campo della sperimentazione a fatica, non era stata identificata la posizione dei provini lungo il cordone di saldatura. Le prove effettuate a supporto dell’attività sui controlli non distruttivi, descritte in seguito, hanno mostrato una possibile diminuzione della resistenza a fatica nella direzione di saldatura. Di conseguenza, lo scatter osservato potrebbe trarre origine anche da questa osservazione, sulla quale sarà condotto un ulteriore specifico approfondimento. I provini realizzati con doppia passata mostrano viceversa un comportamento più omogeneo, con una diminuzione di prestazioni non significativa sia per quelli ricavati da pannelli piani che curvi, nonostante che per i secondi in alcuni casi si sia osservato un leggero undercut. Per questo motivo il carico da applicare in ogni prova è stato valutato sulla base dello spessore effettivamente misurato. I risultati dei provini ricavati da pannelli piani e da pannelli curvi sono perfettamente confrontabili e quindi per maggior chiarezza sono riportati nella fig. 4 con lo stesso simbolo. I risultati relativi al punto finale di chiusura dei cordoni di saldatura mostrano una minore resistenza a fatica, anche a causa di alcune imperfezioni superficiali, descritte in seguito, che saranno oggetto di ulteriori indagini. Le rotture per fatica per i due tipi di giunti, singola e doppia passata, si sono tipicamente nucleate nella zona termicamente alterata in corrispondenza del punto terminale di contatto dello spalla dell’utensile, anche se in questo punto non erano presenti discontinuità geometriche di rilievo. La fig. 5a mostra una tipica rottura per fatica; le superfici di frattura presentano l’aspetto “legnoso” caratteristico di questa lega per cricche in direzione TL, come già osservato in [9] anche per il caso di metallo base. Per quanto concerne l’esatta locazione, le rotture per fatica si sono nucleate indifferentemente sui due lati advancing e retreating del cordone di saldatura, non mostrando quindi un lato più debole, come viceversa osservato in altre sperimentazioni. In due soli casi si sono osservate rotture per fatica nella zona termomeccanicamente alterata, anche se, come mostrato in fig. 5b, prima della rottura finale il difetto si è comunque riunito con altri piccoli difetti di fatica nucleati nella zona termicamente alterata. Per quanto riguarda le chiusure si è osservato in corrispondenza del punto finale di estrazione del pin una traccia, simile a un grippaggio, prodotta dalla filettatura del perno, fig. 6, nella quale in alcuni casi si sono nucleati i difetti per fatica, fig. 7. Questo aspetto sarà ulteriormente indagato su provini fresati in modo da asportare la parte superficiale contenente i difetti.

Generazione di cricche di fatica per la messa a punto delle tecniche di controllo non distruttivo

Aspetto fondamentale per la certificazione di un nuovo processo è l’individuazione di tecniche di controllo non distruttivo in grado di rilevare con elevata confidenza gli eventuali Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno198 2004 - 5 - difetti presenti all’interno di un giunto. Nel caso del Friction Stir Welding Alenia Spazio intende adottare, unitamente al più tradizionale sistema basato sulle correnti indotte, un sistema ad ultrasuoni ad alta definizione, Phased Array, che invece del singolo trasduttore utilizzato nei controlli ultrasonori contiene diversi trasduttori, ognuno dei quali invia un fascio di ultrasuoni ad un determinato angolo e focalizzato ad una prefissata profondità. Il sistema software che coordina i vari elementi elabora le informazioni che da essi provengono fornendo in un'unica passata il quadro completo dei difetti presenti nel componente in esame. Una parte importante dell’attività di collaborazione con Alenia Spazio è stata finalizzata alla definizione e alla caratterizzazione dei difetti di fatica nei cordoni di saldatura. Per ottenere una banca dati sperimentale di difetti realistici su cui tarare i metodi selezionati, quindici pannelli in 2195-T8 saldati sono stati affaticati a diversi livelli. I pannelli utilizzati hanno una larghezza di 200 mm e presentano una saldatura trasversale, come mostrato in fig. 8. Questi pannelli sono stati sottoposti a prova di fatica, Smax=230 MPa, R=0.1; dalla fig. 4 a questa sollecitazione corrisponde una vita media di 109500 cicli. Le prove sono state sospese a numeri cicli compresi fra il 40 e i 60% del numero di cicli a rottura e si sono effettuati i controlli non distruttivi; a tali controlli seguiranno ulteriori affaticamenti e ulteriori controlli, in modo da monitorare la crescita dei difetti. I controlli effettuati su un primo lotto sia con Eddy Current che con Phased Array hanno evidenziato diversi difetti, prevalentemente sul lato advancing del cordone di saldatura. I difetti sono maggiormente presenti nei pannelli estratti dalla parte intermedia e soprattutto dalla parte terminale dei cordone di saldatura (FXX-2 e FXX-3 in fig. 8). Su tali pannelli si sono riscontrate anche rotture precoci, fino al 44% della vita stimata (F27-2, F27-3, F28-2 e F28-3, rotti rispettivamente al 50%, 44%, 55% e 60.5% della vita stimata). Nei pannelli ricavati dalla parte iniziale delle piastre saldate, FXX-1, non sono stati evidenziati difetti neanche nei pannelli affaticati al 55% della vita. Le successive prove effettuate (misura di tensioni residue e propagazione di difetti) descritte in seguito, non hanno mostrato indizi di possibili variazioni della resistenza lungo i cordoni di saldatura. L’unico dato oggettivo resta l’osservazione di “blisters”, piccole aree lucide, assenti nella parte iniziale dei cordoni e progressivamente sempre più presenti lungo le saldature. Come detto precedentemente, su questo punto saranno effettuate ulteriori indagini, principalmente resistenza statica e resistenza a fatica di piccole provette, ricavate in punti diversi di una o più piastre saldate. Come esempio di difettologie, la parte destra della fig. 9 mostra il risultato dell’indagine effettuata con Eddy Current su un campione contenente un piccolo difetto, avente una dimensione superficiale di 2.5 mm e una profondità di 1 mm. Successivamente il campione è stato rotto staticamente per mostrare il difetto presente. Il difetto è stato chiaramente identificato anche con il metodo delle Phased Array, anche se la tecnica non è ancora messa a punto in modo definitivo.

Tensioni residue presenti nei giunti saldati

Nello svolgimento dell’intero lavoro si intende effettuare una misura comparativa delle tensioni residue misurate con estensimetri e con birifrangenza acustica, una tecnica di tipo non distruttivo. Ad oggi è stata effettuata solo una misura, su un primo campione, con estensimetri elettrici. Alcune informazioni riguardanti la metodologia di misura sono riportate in [9]. Si è ipotizzata la simmetria delle tensioni residue sui due lati (advancing e retreating) del cordone, per cui solo metà piastra è stata strumentata con estensimetri. I risultati ottenuti sono mostrati in fig. 10; la tensione residua longitudinale, la più importante dal punto di vista pratico, presenta un picco di circa 130 MPa al centro del cordone di saldatura. Tale valore si mantiene quasi costante nell’area interessata dal contatto della spalla dell’utensile. Oltre questo punto la tensione diminuisce bruscamente fino ad una inversione di segno a circa 21 mm dall’asse del cordone. La tensione residua agente nella direzione del cordone è stata misurata anche in otto diverse stazioni nella direzione longitudinale, per verificare l’eventuale presenza di anomalie che potessero giustificare la minor resistenza a fatica apparentemente osservata nei pannelli ricavati dal tratto terminale del cordone di saldatura, di cui si è trattato nel paragrafo precedente. I risultati ottenuti, riportati in fig. 11, non Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno199 2004 - 6 - mostrano anomalie sull’asse del cordone di saldatura, mentre la stessa misura effettuata sul lato advancing a 12 mm dall’asse del cordone sono di difficile interpretazione. La tensione residua longitudinale sembra comunque diminuire, segno di un minor sovra-riscaldamento, e non aumentare nella direzione del cordone di saldatura. Interessanti per l’interpretazione del fenomeno saranno i risultati della misura effettuata con la birifrangenza acustica, che è una misura a pieno campo e non una misura puntuale come viceversa è la misura estensimetrica.

Propagazione di difetti

Nella attività precedentemente svolta, [9], è stata effettuata anche la caratterizzazione del materiale base, 2195-T8, dal punto di vista della propagazione dei difetti. Per tale attività sono state utilizzate piastre aventi uno spessore di 7 mm; questi risultati, riportati nelle figure 12 e 13, saranno utilizzati per confrontare i risultati relativi al materiale saldato, essendo lo spessore delle piastre oggetto della presente indagine, pari a 5 mm, sostanzialmente simile a quello delle piastre precedentemente analizzate. La velocità di propagazione di un difetto in funzione della variazione nel ciclo di carico del fattore di intensità degli sforzi è stata misurata per due diversi valori del rapporto di sollecitazione R=0.1, fig. 12 e R=0.5, fig. 13, e per i due possibili orientamenti delle fibre, LT e TL. Le differenze nella velocità di propagazione di un difetto nelle due direzioni delle fibre sono molto basse per cui è possibile approssimare tutti i risultati ottenuti con una sola curva media, riportata nelle due figure con segno marcato. Solo nella fase finale della curva di propagazione si osserva un leggero aumento della velocità di propagazione nelle direzione TL rispetto alla direzione LT. Nelle stesse figure sono riportate, come elemento di confronto, anche le curve di propagazione del 2024-T351, tratte da AFGROW, [11], utilizzando fra i diversi modelli di propagazione quello proposto da Harter. Fra le leghe di alluminio il 2024-T351 è una delle leghe nelle quali i difetti propagano più lentamente; in effetti dal confronto si osserva una maggior velocità di propagazione nel 2195-T8, particolarmente per i valori più bassi di DK. Anche il 2024-T351 presenta nella curva di propagazione l’anomalia osservabile nel 2195-T8, caratterizzata da un forte aumento della velocità di propagazione per valori intermedi di DK. Le sole curve medie di propagazione del 2195-T8 sono nuovamente mostrate in fig. 14 unitamente alle rette di best fit ottenute con la legge di Walker, [12] nell’intervallo dei valori medio-alti di DK (superiori a 9 MPa m ). Per il caso in esame si è trovata la relazione:

da = 1.85*10-7 (DK)2.686 (1- R)-0.98 dN Questa modellizzazione sarà utilizzata in seguito per interpretare i risultati relativi alle piastre saldate. Per quanto concerne le piastre saldate, nel caso di difetto che si propaga lungo il cordone di saldatura, si è introdotto il difetto iniziale nel punto in cui termina il contatto con la spalla dell’utensile, dal lato retreating. La propagazione dei difetti è avvenuta in modo regolare, senza scostamenti dalla posizione iniziale. I risultati ottenuti sono riportati nelle fig. 15 e 16 per i due diversi valori del rapporto di sollecitazione esaminati, R=0.1 e R=0.5. Per ciascuna configurazione sono state effettuate due prove nominalmente identiche e volutamente i campioni scelti sono stati il primo e l’ultimo di diverse piastre saldate, vedi fig. 8, per evidenziare eventuali differenze di proprietà meccaniche nella direzione dei cordoni di saldatura. I risultati delle prove non hanno comunque mostrato differenze di rilievo. Per il caso R=0.1 si osserva un aumento della velocità di propagazione rispetto al materiale base, mentre per il caso R=0.5 i due risultati sono praticamente coincidenti. Globalmente si può concludere che la FSW nella lega 2195-T8 non ha un grande effetto sulla velocità di propagazione di un difetto, nel caso in cui questo si propaga nella direzione del cordone stesso. La fig. 17 mostra i risultati relativi al caso di difetto che si propaga nella direzione perpendicolare al cordone di saldatura. In questo caso si osserva un forte aumento della velocità di propagazione di un difetto rispetto al caso del metallo base, anche se il difetto si Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno200 2004 - 7 - propaga verso zone in cui è sempre meno influente l’effetto del riscaldamento prodotto dalla saldatura. Solo ai valori più elevati di DK i risultati tendono a coincidere. Quanto osservato è evidentemente legato alle tensioni residue presenti nei giunti saldati, che aumentano la velocità di propagazione di un difetto fino a che questo non raggiunge una dimensione tale da far rilassare le tensioni interne. Anche in questo caso non si osservano differenze significative nei risultati ottenuti con i due pannelli, uno ricavato nella parte iniziale, e uno nella parte finale, di un cordone di saldatura, come mostrato nel riquadro contenuto in fig. 17. Per tenere conto dell’effetto delle tensioni residue sulla velocità di propagazione di un difetto si è utilizzata una relazione dovuta a Tada e Paris, che fornisce il valore del fattore di intensità degli sforzi “residuo”, cioè prodotto dal campo di tensioni residue, in funzione della dimensione del difetto, noto il valore della tensione residua longitudinale al centro del cordone, s0 e la distanza dall’asse del cordone del punto nel quale la tensione residua cambia di segno, “c”.

1/2 ïì 1+ (a/c) 4 - (a/c) 2 ïü K RES = s o pa í ý 1 (a/c) 4 îï + þï

Sommando il valore di KRES così determinato ai valori massimi e minimi del fattore di intensità degli sforzi nel ciclo di carico, si determinano i valori minimo e massimo “effettivi” del fattore di intensità degli sforzi e quindi il valore effettivo del rapporto di sollecitazione, R. Utilizzando la relazione analitica precedentemente determinata per valutare l’accrescimento di un difetto in funzione di DK e R, (legge di Walker, fig. 14) si è effettuata la previsione dell’accrescimento del difetto nella piastra F69-2 (condizioni d prova: Smax=56.6 MPa, R=0.1). Per la simulazione si è assunto un difetto iniziale avente una semilunghezza a=10 mm, in modo da contenere tutta la simulazione nel tratto lineare della curva di propagazione (DK>9 in fig. 14). Per quanto concerne le tensioni residue, le dimensioni di questa piastra non sono sufficienti per poter sviluppare tutto il campo di tensioni, come deducibile anche dalla fig. 10, per una piastra avente semilarghezza di 100 mm. Per semplicità si è effettuata la misura con un solo estensimetro della tensione agente al centro del cordone di saldatura, che ha fornito un valore di 86.7 MPa e si è assunto che anche nella piastra avente una larghezza di 200 mm la dimensione “c” assumesse il valore c=21 mm. La fig. 18 mostra la schematizzazione utilizzata dal modello di Tada e Paris del campo di tensioni residue (definito dalle due grandezze s0 e “c”, precedentemente descritte) per il caso di piastra non fessurata, mentre la fig. 19 mostra l’andamento del fattore di intensità degli sforzi conseguente allo stato di tensioni residue, in funzione della semi-dimensione del difetto. Il fattore di intensità degli sforzi residue aumenta con la dimensione del difetto fino a raggiungere un valore massimo, di circa 14 MPa m per una dimensione di difetto a=13 mm per poi diminuire a causa della ridistribuzione delle tensioni interne. La fig. 20 mostra il risultato della simulazione, effettuata nel modo precedentemente descritto, sulla piastra F69- 2, confrontati con i risultati della prova. Nella stessa figura sono riportati anche i risultati di una simulazione effettuata con il solo metallo base, cioè senza tenere conto della presenza di tensioni residue. I risultati mostrano l’importanza delle tensioni residue sulla propagazione di un difetto, che è comunque perfettamente prevedibile nota la distribuzione delle tensioni residue unitamente alle proprietà del metallo base.

CONCLUSIONI

Le problematiche di fatica, Damage Tolerance e tenso-corrosione in giunzioni realizzate con Friction Stir Welding nella lega alluminio-litio 2195-T8 costituiscono l’oggetto di una ricerca svolta congiuntamente dal Dipartimento di Ingegneria Aerospaziale dell’Università di Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno201 2004 - 8 -

Pisa e la Alenia Spazio. Fra gli scopi della ricerca vi è anche la caratterizzazione delle proprietà meccaniche dei giunti saldati, finalizzata alla qualifica del processo di saldatura. Allo stato attuale circa l’80% dell’attività è stata svolta; sono in fase di allestimento le prove di suscettibilità a tenso-corrosione e le prove di resistenza a frattura che, unitamente ad alcuni approfondimenti di temi già sviluppati, concluderanno la ricerca. La lega 2195-T8 presenta delle ottime proprietà meccaniche ed anche le giunzioni Friction Stir Welding possiedono caratteristiche elevate. Buone proprietà sono state rilevate anche nelle giunzioni saldate con una doppia passata; tale operazione si rende necessaria nelle saldature circonferenziale dei recipienti pressurizzati, in modo da chiudere la saldatura stessa. Anche il punto nel quale la saldatura viene interrotta è stato oggetto di indagine. Dal lavoro fin ad ora svolto è possibile trarre le seguenti conclusioni: · Tensione di snervamento e tensione di rottura dei giunti saldati sono rispettivamente pari al 50% e al 70% delle stesse proprietà misurate nel metallo base. L’allungamento a rottura è più alto di quello del metallo base, ma la deformazione plastica riguarda solo il cordone di saldatura, fornendo così l’impressione di una rottura fragile. · La durezza nel cordone di saldatura è inferiore del 37% rispetto alla durezza del metallo base. Il profilo di microdurezza si presenta simmetrico, indice di un buon processo di saldatura. · Una seconda passata di saldatura aumenta le dimensioni del nugget, rendendo più brusco il passaggio tra la zona termomeccanicamente alterata e il metallo base. Ancora nei cordoni con doppia passata, analisi micrografiche hanno mostrato una grana del nugget centrale più fine, con una maggiore dispersione degli ossidi rispetto al cordone a singola passata. · La resistenza a fatica dei giunti saldati è molto elevata; le tensioni massime ammissibili nella zona degli alti numeri di cicli sono dell’ordine di 180 MPa (R=0.1). Anche la dispersione dei risultati è elevata; tale dispersione potrebbe essere causata da una diminuzione sistematica della resistenza a fatica lungo i giunti saldati. Questo aspetto sarà verificato con ulteriori prove specifiche. Le rotture per fatica si sono tipicamente nucleate nella zona termicamente alterata, indifferentemente dai due lati advancing e retreating. · E’ stata effettuata una misura con estensimetri delle tensioni residue introdotte dal processo di saldatura. La tensione residua agente nella direzione del cordone di saldatura, la più importante dal punto di vista pratico, presenta un picco di circa 130 MPa al centro del cordone di saldatura. Tale valore si mantiene quasi costante nell’area interessata dal contatto della spalla dell’utensile. Oltre questo punto la tensione diminuisce bruscamente fino ad una inversione di segno a circa 21 mm dall’asse del cordone. · La velocità di propagazione di un difetto lungo il cordone di saldatura è leggermente più alta rispetto alla velocità di propagazione di un difetto nel metallo base, per un rapporto di sollecitazione R=0.1. Nel caso R=0.5 i due risultati sono praticamente coincidenti. · Nel caso di difetto disposto perpendicolarmente al cordone di saldatura si osserva un forte aumento della velocità di propagazione rispetto al caso del metallo base. Tenendo opportunamente conto delle tensioni residue, attraverso un “fattore di intensità degli sforzi residuo”, si è ottenuto un ottimo accordo fra i risultati sperimentali e la previsione ottenuta utilizzando le curve di propagazione del metallo base e valutando durante la simulazione la variazione del valore effettivo del rapporto di sollecitazione.

BIBLIOGRAFIA

[1] M. Braukus, J. Malone "Shuttle Super Lightweight Fuel Tank Completes Test Series". NASA Press realise July 1996. [2] W.M. Thomas, E.D. Nicholas, J.C. Needham, M.G. Church, P. Templesmith, C.J. Dawes (1991). International Patent Application No. PCT/IGB92/02203 and GB Patent Application No. 9125978.9. [3] R. Talwar, D. Bolser, R. Lederich, J. Baumann "Friction stir welding of airframe structures". Proceedings of the 2nd International Symposium on Friction Stir Welding, Gothenburg, Sweden, 2001. Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno202 2004 - 9 -

[4] D. Waldron "Application of Friction Stir Welding for Delta Rocket Fuel Tanks". Proceedings of the 1st International Symposium on Friction Stir Welding, Thousand Oaks, California, USA, 1999. [5] I. Henderson "Exploiting friction stir welding in explosively formed aluminium boat hull construction". Proc. Conf. 'INALCO' 98. 7th International Conference on Joints in Aluminium. Cambridge, UK, 1998, pp. 251-257. [6] T. Kawasaki, T. Makino, S. Todori, H. Takai, M. Ezumi, Y. Ina "Application of friction stir welding to the manufacturing of next generation "A-train" type rolling stock". Proc. 2nd International Symposium on Friction Stir Welding, Gothenburg, Sweden, 2000. [7] M.R. Johnsen "Friction stir welding takes off at Boeing". Welding Journal 78 (2), 1999, pp. 35-39. [8] A. Strombeck, C. Schilling, J.F. dos Santos "Robotic friction stir welding -Tool, Technology and Applications". 2nd International Symposium on Friction Stir Welding, Gothenburg, Sweden, 2000. [9] A. Lanciotti, C. Polese "Alcune proprietà di giunzioni 'Friction Stir Welding' in due leghe di alluminio di impiego aerospaziale" Memoria presentata al Convegno "Friction Stir Welding per Applicazioni Aerospaziali"; Pisa 1 Aprile 2003. [10] C. Dalle Donne “Proprietà e caratterizzazione di laboratorio del FSW di leghe di alluminio aerospaziali”. Memoria presentata al Convegno “Applicazioni e Prospettive del Friction Stir Welding”, Cadriano (BO), 11 Ottobre, 2001. [11] J.A. Harter “AFGROW user guide and technical manual”. Air Force Research Laboratory, VA-WP-TR-1999-3016, 1999 [12] K. Walker "The Effect of Stress Ratio during Crack Propagation and Fatigue for 2024-T3 and 7075-T6 Aluminium", in "Effects of Environment and Complex Load History on Fatigue Life", ASTM STP 462, pp. 1-14, 1970. [13] H. Tada, P.C. Paris “The stress intensity factor for a crack perpendicular to the welding bead”. International Journal of Fracture, n. 21, pp.279-284, 1983.

Ringraziamenti

Gli autori ringraziano il laureando Federico Ferrari per l'aiuto fornito nell’esecuzione delle prove sperimentali e l’elaborazione dei risultati ottenuti. Per lo svolgimento di questa attività il Dipartimento ha usufruito di un finanziamento del Ministero dell’Istruzione, dell’Università e della Ricerca, programma COFIN 2003/2004.

Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno 2004 - 10 - 203

Materiale base A% 700 FSW 12

600 10

500 ] 8

MPa 400

S [ 6

300 4

200 2

100 0

0

Sy Su A%

Fig. 1 - Proprietà meccaniche del metallo base e dei giunti saldati.

180 Singola passata Doppia passata 160 materiale base

140

120 HV0.1

100

80 Retreating Advancing

-16 -12 -8 -4 0 4 8 12 16 Weld bead x [mm]

Fig. 2 - Microdurezze nei cordoni saldati (singola e doppia passata). Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno 2004 - 11 - 204

a) singola passata 5

b) doppia passata 5

Fig. 3 - Macrografie di cordoni saldati (singola e doppia passata).

350 R=0.1 Singola passata Doppia passata 300 Chiusura foro Materiale base

250 [MPa] doppia passata max materiale base S singola passata 200

chiusura foro

150 105 106 Cicli

Fig. 4 - Risultati delle prove di fatica. Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno 2004 - 12 - 205

b)

Zone preferenziali di innesco delle rotture per fatica

a) 1 mm

0.5 mm

Fig. 5 - Tipiche rotture per fatica nei giunti saldati.

1 mm

0.2 mm

0.2 mm

Fig. 6 - Dettaglio di un punto di chiusura. Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno 2004 - 13 - 206 2 mm

0.1 mm

Fig. 7 - Innesco di una rottura per fatica in un punto di chiusura.

Punti di innesco

FXX-1 FXX-2 FXX-3 Fig. 8 - Dimensioni e numerazione dei pannelli con saldatura trasversale.

450 Inizio saldatura

200

2.5 mm

Fig. 9 - Cricca di fatica e sua rilevazione con Eddy Current. Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno 2004 - 14 - 207

140

120

100 y Lato Retreating 80 Sres x 60

40

S res [MPa] 20

0

-20

-40 0 50 100 150 200 250 x [mm] Fig. 10 - Tensioni residue longitudinali misurate nella sezione centrale di una piastra saldata.

180

160

140

120 Sres [MPa] 100

80 S y

60 res x 40

20 x=-12 mm (lato advancing) Asse del cordone 0 -400 -300 -200 -100 0 100 200 300 400 Inizio saldatura y [mm]

Fig. 11 - Tensioni residue longitudinali misurate in due sezioni lungo un cordone di saldatura. Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno 2004 - 15 - 208

0.01 2195-T8, Materiale base; R=0.1; L-T; T-L; t = 7 mm

2024-T351 1E-3

1E-4 TL LT da/dN [mm/ciclo]

1E-5

1E-6 2 3 4 5 6 7 8 9 10 20 30 40 50 DK [MPa*m0.5]

Fig. 12 - Velocità di propagazione di un difetto nel 2195-T8, R=0.1.

0.01

2024-T351

1E-3

LT 1E-4 TL

da/dN [mm/ciclo]

1E-5

2195-T8, Materiale base; R=0.5; L-T, T-L; t = 7 mm 1E-6 2 3 4 5 6 7 8 9 20 30 40 50 0.5 DK [MPa*m ]

Fig. 13 - Velocità di propagazione di un difetto nel 2195-T8, R=0.5. Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno 2004 - 16 - 209

0.01 da/dN=1.85*10-7(DK)2.686(1-R)-0.98

R=0.1 1E-3 R=0.5

1E-4 da/dN [mm/ciclo]

1E-5

2195-T8, Materiale base; t = 7 mm

1E-6 2 3 4 5 6 7 8 9 10 20 30 40 50

DK [MPa*m0.5] Fig. 14 - Approssimazione delle curve di propagazione con la legge di Walker.

0.01 F68-1 FSW; Smax=71.9 MPa F68-3 FSW: Smax=73.6 MPa 2195-T8 1E-3 R=0.1

1E-4 da/dN [mm/ciclo] 1E-5

1E-6 2 3 4 5 6 7 8 9 10 20 30 40 50 DK [MPa*m0.5]

Fig. 15 - Propagazione di difetti nella direzione del cordone di saldatura; R=0.1. Giornata di Studio: “Friction Stir Welding Day", Torino, 15 Giugno 2004 - 17 - 210

0.01 F70-3 FSW; Smax=73 MPa F70-1 FSW; Smax=73 MPa 2195-T8 1E-3 R=0.5

1E-4

da/dN [mm/ciclo]

1E-5

1E-6 2 3 4 5 6 7 8 9 10 20 30 40 50 DK [MPa*m0.5] Fig. 16 - Propagazione di difetti nella direzione del cordone di saldatura; R=0.5.

0.01 2195-T8 (LT, TL) F69-1 FSW: Smax=74.8 MPa F69-2 FSW; Smax=56.6 MPa

1E-3 R=0.1 200 FXX-2 1E-4 450 da/dN [mm/ciclo]

1E-5 FXX-1 Inizio sald.

1E-6 2 3 4 5 6 7 8 9 10 20 30 40 50 DK [MPa*m0.5]

Fig. 17 - Propagazione di difetti nella direzione perpendicolare al cordone di saldatura; R=0.1. Giornata di Studio: “Friction Stir Welding Day", Torino, 15211 Giugno 2004 - 18 -

100 y 14

σο Sres 12 80 x 2a 10 60 ) 2 1− ξ 0.5 8 40 σ = σ 0 1+ ξ 4

(MPa) ξ = x/c 6 (MPa*m res 20 S σ = 86.7 MPa

0 RES c = 21 mm K 4 0

c 2 -20 0 0 102030405060 0 102030405060 x (mm) a (mm) Fig. 18 - Schematizzazione della tensione Fig. 19 - Fattore di intensità degli sforzi residua longitudinale. prodotto dalle tensioni residue.

50

40

30

a (mm) 20

F69-2 FSW (R=0.1) 10 Mat. base (R=0.1) Mat. base+Tens. residue

0 0 50000 100000 150000 Cicli Fig. 20 - Propagazione di un difetto nella direzione perpendicolare al cordone di saldatura; risultati sperimentali e previsioni.

212

100 6013-T4 FSW 2219-T851 FSW 80

60

40

20 Retreating side Advancing side FSWeld 0

-20

-40 Longitudinal residual stress [MPa]

-100 -50 0 50 100

Di t f th t li f th ld [ ] Fig.15.- Residual stresses.

213

XVIII CONGRESSO NAZIONALE AIDAA 19-22 settembre 2005 VOLTERRA (PI)

PROPRIETÀ MECCANICHE DI GIUNZIONI "FRICTION STIR WELDING" IN UNA LEGA ALLUMINIO-LITIO PER APPLICAZIONI SPAZIALI

A. LANCIOTTI1, R. LAUDATO2, C. POLESE1

1Dipartimento di Ingegneria Aerospaziale, Università di Pisa, Pisa 2Alenia Spazio S.p.A., Torino

SOMMARIO

Nel presente lavoro si descrivono i risultati di una serie di prove sperimentali per la caratterizzazione delle proprietà generali di giunzioni realizzate con la tecnologia "Friction Stir Welding" nella lega alluminio-litio 2195-T8P4, impiegando la particolare tecnica del "Retractable Pin Tool". L’attività è stata svolta nell’ambito di una collaborazione fra il Dipartimento di Ingegneria Aerospaziale dell’Università di Pisa e Alenia Spazio, ed è mirata alla qualifica del processo di saldatura all’interno del programma FAST 2. Gli aspetti investigati hanno riguardano l’analisi della microstruttura del giunto, la resistenza statica, la micro-durezza, la resistenza a tenso- corrosione e la resistenza a fatica. I risultati delle prove, confrontati con gli stessi risultati ottenuti dal metallo base, hanno mostrato le ottime proprietà dei giunti saldati. E’ stato inoltre messo a punto un sistema di un controllo non distruttivo per l’ispezione dei cordoni di saldatura.

1. INTRODUZIONE

I requisiti di progetto nel campo delle strutture di impiego spaziale sono sempre state la leggerezza della struttura associata ad ermeticità ed elevata resistenza meccanica. Nei più recenti sviluppi nell’industria del settore hanno trovato largo impiego le leghe alluminio-litio di nuova generazione. Tali leghe, caratterizzate rispetto alle precedenti da una ridotta percentuale di litio come elemento allegante, consentono comunque consistenti risparmi ponderali grazie alla diminuzione di peso specifico ed alle elevate proprietà meccaniche. Queste leghe essendo anche saldabili ed impiegabili a temperature criogeniche sono state impiegate per la costruzione di serbatoi per combustibile di veicoli spaziali in sostituzione di leghe più tradizionali, come la 2219. In queste applicazioni solo la saldatura riesce a soddisfare i requisiti di ermeticità della struttura. Per tale motivo la saldatura in ambito spaziale è sempre stato ritenuta un processo strategico e notevoli sforzi tecnici e tecnologici sono stati intrapresi con lo scopo di ottimizzare la qualità e le prestazioni dei giunti realizzati. Fra le varie tecniche di saldatura particolarmente interessante appare la Friction Stir Welding (FSW), una tecnica relativamente recente sviluppata dal The Welding Institute (TWI, Cambridge, UK) [1] con la quale è possibile realizzare giunzioni saldate senza raggiungere la temperatura di fusione del materiale. La giunzione si realizza tramite il calore sviluppato per attrito da un perno che, ruotando, avanza lungo l’interfaccia fra le due lamiere accostate di testa. Il processo è molto simile ad una estrusione "in situ", che sposta il metallo dalla parte anteriore a quella posteriore dell'utensile. Il minor apporto termico che compete a questa tecnica migliora notevolmente le proprietà meccaniche dei giunti saldati, mentre la minore temperatura diminuisce la possibilità di contaminazione del metallo. Allo stato attuale la saldatura FSW di leghe di alluminio è diventata una realtà industriale. In un’ottica di continua innovazione Alenia Spazio ha individuato la FSW come "key tecnology" ed ha acquisito un impianto specifico dedicato alla FSW, sviluppando una attività mirata alla ottimizzazione dei parametri di processo per alcuni materiali di interesse aerospaziale. Tale attività si è concretizzata con la costruzione di un serbatoio per ossigeno liquido (LOX), all'interno del programma FAST 2 (Future Advanced Space Transportation Technologies). In questo ambito si introduce un’attività di ricerca congiunta con il Dipartimento di Ingegneria Aerospaziale dell’Università di Pisa su giunzioni saldate nella lega alluminio-litio 2195-T8P4. L'attività è finalizzata alla qualifica del processo di saldatura, requisito necessario per la produzione di moduli spaziali e serbatoi criogenici, come possibile alternativa ai processi TIG e plasma attualmente impiegati. 214

Un punto debole della tecnologia FSW è costituito dalla presenza nel giunto del foro lasciato dall’utensile in fase di estrazione; questo "difetto" non è ammissibile nella realizzazione di saldature circonferenziali su vessel di impiego spaziale, mentre in saldature lineari il difetto può essere traslato su un tallone terminale, che viene successivamente rimosso. Per questo motivo Alenia Spazio si avvale della nuova tecnologia "Retractable Pin Tool" che permette di realizzare saldature su traiettorie chiuse, in quanto la saldatura si conclude con un tratto in sovrapposizione con il tratto iniziale, durante il quale il perno dell'utensile viene progressivamente retratto. Nel presente lavoro si descrivono gli aspetti fondamentali del processo di saldatura unitamente ai risultati di una serie di prove di caratterizzazione di giunzioni saldate in lega di alluminio-litio 2195-T8P4. Gli aspetti investigati riguardano le caratteristiche microstrutturali del giunto e le principali proprietà meccaniche come la resistenza statica, la resistenza a fatica e la resistenza a tenso-corrosione. Parallelamente è stata valutata l'adeguatezza di diverse tecniche di controllo non distruttivo nell'individuazione dei difetti peculiari di questo tipo di saldatura, attività anch'essa necessaria per la qualifica del processo. I risultati delle prove, confrontati con gli stessi risultati ottenuti sul metallo base, hanno mostrato le ottime proprietà dei giunti saldati.

1.1 Il serbatoio criogenico del programma FAST 2 di Alenia Spazio Un’ampia attività di ricerca ha preso il via all’interno del programma di Alenia Spazio FAST 2 finalizzata alla realizzazione del serbatoio per l’ossigeno liquido le cui dimensioni finali, indicate in fig. 1, sono di una capsula di diametro 1.3 m per un’altezza totale di 2.25 m. I requisiti di progetto prevedono una struttura ermetica di minimo peso in grado di sostenere le sollecitazioni meccaniche dovute alla massima pressione di esercizio prevista (4+0.4 bar) e alle forti accelerazioni radiali, 3.5 g, e assiali, 2.5 g, a cui è soggetta la struttura, con temperature di esercizio previste variabili tra la temperatura ambiente, 20°C, e la temperatura del propellente LOX, -196°C. La soluzione costruttiva adottata da Alenia Spazio consiste in due calotte dotate di flangia di interfaccia con il lanciatore e un cilindro centrale ottenuto dall’unione, mediante giunti di testa, di quattro pannelli con estensione 90°. Le calotte sono state progettate da Alenia Spazio e realizzate dalla società americana Spincraft impiegando il processo di "spin forming". I pannelli, progettati e realizzati da Alenia Spazio, sono stati realizzati tramite lavorazione alle macchine e successivamente curvati mediante il processo di "brake forming". Tutti i sub-componenti, con uno spessore minimo di 5 mm, sono poi stati sottoposti ad un processo di invecchiamento e trattamento superficiale con Alodine 1200, [2]. Nell’ambito di questo avanzato programma tecnologico si sono valutate le prestazioni di materiali e processi tecnologici innovativi che possono essere efficacemente impiegati per tale applicazione.

1.2 Scelta del materiale: Al-Li 2195-T8P4 Nell’industria aerospaziale il processo di continua ricerca di materiali di elevate caratteristiche meccaniche specifiche ha individuato le elevate potenzialità del litio che, introdotto in ridotta percentuale (massimo 3%) come elemento allegante nelle leghe leggere, consente rispetto alle leghe leggere tradizionali un aumento di circa il 50% della resistenza statica e del 5% del modulo elastico ed una riduzione del peso specifico del 5%. Nella progettazione di un componente spaziale si devono considerare anche altri fattori come le forti sollecitazioni di natura termica, determinate dalle rapide escursioni in temperatura e al bombardamento continuo in orbita di pulviscolo e radiazioni solari, e l’esposizione ad un ambiente fortemente aggressivo, come ad es. il contatto con propellenti. Le nuove leghe alluminio-litio hanno dimostrato di mantenere stabili le elevate proprietà meccaniche fino a valori estremi della temperatura ed anche una notevole resistenza a corrosione e tenso-corrosione. Pertanto, essendo anche saldabili, queste leghe di nuova generazione rispondono perfettamente ai requisiti di progetto di leggerezza, alta resistenza ed ermeticità della struttura tipici delle strutture di impiego spaziale e possono essere efficacemente impiegate per la costruzione di serbatoi per combustibile di veicoli spaziali in sostituzione di leghe più tradizionali, come la 2219. Per la presente attività Alenia Spazio ha selezionato la lega Al-Li 2195-T8P4, commercializzata dalla Reynolds Metals con il nome commerciale di WeldaliteTM. La lega Al-Li 2195 è una lega caratterizzata da un’elevata resistenza statica e da un basso peso specifico. Le caratteristiche chimiche della lega sono riportate in tabella 1.1.

Lega Al Cu Li Si Fe Mn Mg Zn Ti AG 2195 94.8 4 1 0.03 0.05 0.05 0.4 0.01 0.02 0.4 Tabella 1.1 215

Il trattamento termico effettuato (T8P4) consiste in una tempra di solubilizzazione seguita da deformazione a freddo e infine da un invecchiamento artificiale. Il trattamento di tipo T8 è impiegato generalmente per migliorare la resistenza a corrosione intergranulare nelle leghe della serie 2000 ed è facilmente realizzabile su lamiere e piastre.

1.3 Scelta della tecnologia di giunzione: il Friction Stir Welding Le tecniche di saldatura sono state largamente impiegate nelle giunzioni di strutture spaziali. Comunque sussistono tuttora problemi strettamente connessi con le proprietà del cordone di saldatura e la possibile presenza di difetti. Di conseguenza nel settore si svolgono continuamente ricerche per sviluppare nuove tecniche di saldatura o per migliorare quelle già esistenti. Per molti anni la saldatura TIG (Tungsten Inert Gas) è stata la tecnologia leader per le leghe di alluminio. Successivamente è stata progressivamente sostituita dalla saldatura VPPA (Variable Polarity Plasm Arc Welding) che è attivamente impiegata per la costruzione di importanti strutture aerospaziali. Fra le tecnologie di saldatura di recente introduzione quella che appare più promettente è la cosiddetta tecnica di saldatura "solida", Friction Stir Welding (FSW), brevettata nel 1991 dal The Welding Institute (TWI, Cambridge, UK) con la quale è possibile realizzare giunzioni saldate senza portare a fusione il materiale. Questa consente di realizzare giunzioni anche nelle leghe aerospaziali ad alta resistenza delle serie 2000 e 7000 che risultano difficili da saldare con i processi tradizionali di saldatura per fusione. Un giunto di testa tramite il processo Friction Stir si realizza inserendo un utensile cilindrico posto in rotazione all’interfaccia delle due lamiere da unire. Le piastre devono essere riferite rigidamente su una piastra di fissaggio in modo da prevenire uno scostamento relativo dei lembi da unire. Uno schema di un utensile generico è riportato in fig. 2. L’utensile consiste in uno spallamento con un pin profilato che emerge da esso. Per iniziare il processo di saldatura l’utensile posto in rotazione viene progressivamente inserito tra le due lamiere da unire finchè una parte sostanziale dello spallamento non entra in contatto con la superficie del pezzo al fine di evitare l’espulsione del materiale malleabilizzato. Per produrre una saldatura con completa penetrazione nello spessore l’utensile deve essere progettato in modo tale che l’estremità del pin sia vicino alla superficie inferiore del giunto quando lo spallamento è in contatto con le superfici. Quindi giunzioni di pezzi con vari spessori richiedono o utensili con diverse lunghezze del pin o un utensile speciale con pin di lunghezza variabile o continuamente retrattile. Una volta che la penetrazione è completa, la traslazione dell’utensile lungo la linea di giunzione consente la realizzazione effettiva della saldatura; a causa dell’attrito e della pressione dovute al movimento relativo, i lembi da saldare si surriscaldano e diventano plastici senza raggiungere il punto di fusione del materiale. In questo modo il materiale plasticizzato viene estruso attorno al pin e forgiato dal bordo di uscita dello spallamento. Il processo può essere visualizzato come un’estrusione "in-situ". Lo stampo di estrusione lungo la lunghezza della saldatura è composto dalla piastra di sostegno, dallo spallamento e dal pin dell’utensile e dal materiale più freddo del pezzo. L’asse longitudinale dell’utensile si può definire come l’asse di forgiatura e la direzione di traslazione dell’utensile come l’asse di estrusione. I vari tipi di utensile con diverse geometrie riescono a realizzare uno stato di alta pressione idrostatica lungo la linea del giunto, causando il consolidamento del materiale plasticizzato. Alenia Spazio ha deciso di introdurre il processo FSW nei propri processi produttivi e di impiegarlo in questa applicazione per la saldatura della lega alluminio-litio 2195 nello spessore di 5 mm. La particolare tecnica impiegata in Alenia Spazio si avvale di uno speciale tipo di utensile noto come "Retractable Pin Tool". Questo moderno sistema di pin, fig. 3, è stato realizzato da J. Ding presso il Marshall Space Flight Center e brevettato dalla Nasa, [3]. Il sistema è strutturato come un normale utensile FSW ma il pin ha la possibilità di essere estratto dal giunto all’interno della spalla elettricamente, idraulicamente o pneumaticamente; la spalla e il pin sono completamente indipendenti perché gestiti da due diversi sistemi di attuazione. Con questo utensile si riesce ad evitare la comparsa in fase di estrazione del "Key-hole" che caratterizza le saldature Friction realizzate con utensile monolitico: il "key-hole" infatti è il foro che l’utensile Friction lascia in fase di estrazione dal giunto saldato. In fig. 4 viene presentato un confronto tra un punto di chiusura di un cordone Friction Stir Welding realizzato con la tecnologia del "retractable pin tool" e un "key-hole". Mentre nella realizzazione di giunzioni di tipo lineare la presenza del foro al termine del cordone di saldatura non comporta problemi strutturali, poiché può essere previsto un opportuno tallone di uscita da rimuovere successivamente, il "retractable pin tool" risulta di fondamentale importanza nella realizzazione di cordoni di saldatura su traiettorie chiuse, dove non è accettabile la presenza del foro finale come nel caso del tank criogenico analizzato. In questa applicazione la tecnica adottata da Alenia Spazio consiste nel realizzare le saldature circonferenziali compiendo un’intera rotazione con l’utensile attorno al componente ripassando sul punto di inizio della saldatura stessa; superato questo punto inizia la fase di estrazione del pin fino alla sua completa retroazione all’interno della spalla. Si avrà quindi necessariamente una porzione del cordone in cui è stata realizzata una seconda passata di saldatura. Il fattore chiave è nel sistema di controllo della spalla, in 216 spostamento o in forza, indipendente dal sistema di attuazione del pin. Ciò consente di mantenere la spalla a contatto con il giunto durante la fase di estrazione del pin e impedire quindi la fuoriuscita del flusso di materiale ancora allo stato pastoso; in fig. 5 si riporta schematicamente la dinamica dell’estrazione del "retractable pin tool". Nella foto successiva, fig. 6, si riporta un particolare della realizzazione di saldature FSW sul tank criogenico; si nota lo scalo di montaggio che sostiene la struttura da saldare e l’utensile Friction. Il moto di avanzamento in questo caso viene conferito allo scalo di montaggio mentre l’utensile è fisso. Per ottenere una giunzione priva di difetti risulta importante la messa a punto dei parametri di saldatura, cioè la velocità di traslazione orizzontale della spalla, la velocità di rotazione dell’utensile e la velocità di rotazione della spalla. La saldatrice impiegata in Alenia Spazio è un impianto a controllo numerico prodotto dalla casa costruttrice svedese ESAB che consente un monitoraggio continuo dei parametri tecnologici della saldatura impostati nel programma iniziale. Le piastre utilizzate nel presente lavoro sono state tutte saldate con un programma di saldatura in controllo spostamento per quanto riguarda il pin e in controllo forza per quanto riguarda la spalla. I valori esatti dei parametri di saldatura impiegati per la realizzazione dei giunti sono proprietà di Alenia Spazio.

1.4 Microstruttura del giunto Friction Stir Welding Le operazione di estrusione e forgiatura, unite con la rotazione dell’utensile, sono responsabili della tipica microstruttura associata ad un processo FSW mostrata in fig. 7a. Le zone caratteristiche del giunto FSW possono essere distinte in due categorie: la zona termo-meccanicamente alterata, TMAZ, e la zona termicamente alterata, HAZ. La HAZ si trova immediatamente a lato del cordone di saldatura ed è simile alla zona termicamente alterata che risulta da una processo di saldatura per fusione convenzionale: il materiale ha subito un ciclo termico che ne ha modificato la microstruttura e/o le proprietà meccaniche. A seconda del tipo di lega impiegata, il trattamento termico iniziale e la vicinanza al centro del cordone il processo da luogo nella zona HAZ del FSW ad ingrossamento o dissoluzione dei precipitati, ricristallizzazione o aumento delle dimensioni del grano. La TMAZ di un giunto FSW può essere considerata analoga alla zona di fusione di una saldatura convenzionale eccetto che, invece di essere fuso, il materiale è deformato plasticamente dall’utensile. All’interno della TMAZ si possono distinguere tre regioni diverse; il "nugget", che ha subito la deformazione plastica più severa, è caratterizzato da una struttura dalla grana fine, relativamente equiassica e ricristallizzata. La larghezza del nugget è tipicamente pari o leggermente maggiore al diametro del pin. Al di fuori del nugget su entrambi i lati c’è una seconda regione che è stata deformata in grado minore e che, a seconda della lega, può mostrare o no segni di ricristallizzazione. Le leghe di alluminio possono essere notevolmente deformate in modo plastico ad alte temperature senza mostrare evidenza di ricristallizzazione e solitamente mostrano un confine marcato tra zona ricristallizzata e zona solo deformata. Nella maggior parte dei materiali invece la zona distinta ricristallizzata, il nugget, è assente e l’intera zona TMAZ sembra essere ricristallizzata. In fig. 7a la demarcazione tra il nugget e il resto della zona TMAZ è abbastanza netta. La deformazione dei grani del metallo base si manifesta come una flessione nel piano della sezione metallografica. L’area immediatamente sotto lo spallamento dell’utensile, che è chiaramente parte della TMAZ, può rappresentare una categoria separata poiché solitamente la struttura del grano qui è diversa. Questa terza sub-zona della TMAZ viene denominata "flow arm". Qui la microstruttura è determinata dallo sfregamento della parte posteriore dello spallamento e dal rapido raffreddamento del materiale. E’ da notare che il processo FSW non è simmetrico rispetto alla mezzeria del giunto: l’"advancing side" della saldatura è definito come il lato in cui il vettore velocità di rotazione dell’utensile ha la stessa direzione del vettore della velocità di traslazione dell’utensile rispetto al pezzo saldato. Il "retreating side" è il lato in cui i vettori sono in direzione opposta. Si vedrà che le proprietà meccaniche possono essere leggermente diverse sui due lati. Il ridotto apporto di calore associato a questo cosiddetto processo di saldatura "solida" diminuisce fortemente le possibilità di contaminazione del metallo. Praticamente si può osservare che alcuni difetti tipici della saldatura classica per fusione, come porosità, cricche da ritiro e segregazioni, scompaiono. Tuttavia queste caratteristiche non fanno si che il materiale saldato abbia le stesse proprietà del materiale base da cui deriva. Sebbene sia stato dimostrato che la FSW sia un processo estremamente robusto in termini di intervalli dei parametri di saldatura che possono produrre giunzioni accettabili, è stato osservato che il cordone di saldatura può contenere alcuni tipi di difetti come discontinuità, microvuoti e microimperfezioni strutturali. E’ stata quindi necessaria una fase di messa a punto dei parametri del processo necessaria ad evitare la presenza di tali difettologie, [4].

217

2. ATTIVITA’ SPERIMENTALE

2.1 Materiale di prova Il materiale iniziale è costituito da una serie di lastre di varie dimensioni nella lega Al-Li 2195-T8P4 in spessore 5 mm ottenute per fresatura meccanica da pannelli di spessore pari a 13 mm. Per le prove di caratterizzazione meccanica Alenia Spazio ha realizzato due serie di campioni saldati con la tecnologia "retractable pin tool". La prima serie è costituita da piastre piane aventi dimensioni finali di 300*850 mm, ottenute saldando piastre di 150*800 mm con il giunto disposto parallelamente alla direzione di laminazione del materiale. La seconda serie è costituita da piastre sia piane che curve (raggio di curvatura di 480 mm) contenenti tre tratti di saldatura con doppia passata intervallati da tre chiusure del cordone. Tutte le prove sono state condotte nella condizione "as welded" tranne poche prove di confronto su provini con cordone rasato tramite fresatura. Nella prima fase dell’attività sono state valutate le proprietà meccaniche generali sia del materiale base che dei giunti FSW.

2.2 Macro e micrografia dei giunti saldati FSW Presso il laboratorio metallurgico di Alenia Spazio sono state eseguite una serie di metallografie nei giunti realizzati sia con singola che con doppia passata. Le macrografie trasversali dei giunti riportate in fig. 7 mostrano chiaramente alcune caratteristiche di una giunzione FSW: il nugget centrale, caratterizzato da una grana fortemente equiassica, la zona termomeccanicamente alterata (TMAZ) e la zona termicamente alterata (HAZ). Sono altresì visibili i cosiddetti "onion rings" al centro del nugget e l’elevata fibrosità del materiale base già osservata nel precedente lavoro, [4]. L’aspetto del cordone risulta visibilmente diverso nei due casi di singola e doppia passata: con una seconda passata si ha una maggiore uniformità del giunto, aumentano le dimensioni del nugget e diventa quindi molto più brusco il passaggio tra TMAZ e metallo base. Inoltre appare una grana del nugget centrale più fine con una maggiore dispersione degli ossidi rispetto al cordone a singola passata. In fig. 7b si riporta un dettaglio micrografico dell’interfaccia tra la zona termo-meccanicamente alterata e il nugget in un giunto a doppia passata, dove si nota infatti la rapida transizione tra la grana fine ed equiassica del nugget e l’aspetto "legnoso" e stratificato del materiale base 2195-T8P4. L’altra micrografia del nugget di una doppia passata, fig. 7b, evidenzia la struttura granulare molto fine in cui si possono distinguere gli ossidi introdotti con il processo di saldatura sotto forma di piccole particelle nere. Analizzando invece la superficie esterna superiore del giunto in condizioni "as welded" si distingue la caratteristica successione dei "feed" semicircolari il cui diametro coincide con la larghezza della spalla dell’utensile; il passo tra un feed e il successivo è diretta conseguenza della composizione del moto di avanzamento della spalla e della rotazione dell’utensile. L’aspetto dei feed, che nella parte iniziale del cordone è generalmente molto netta e priva di difetti, fig. 8, comincia a presentare nella parte centrale e soprattutto nella parte finale della saldatura delle vescicolature, "blister", che costituiscono residui di materiale sul cordone di aspetto brillante e tondeggiante dovuti probabilmente al riscaldamento della spalla.

2.3 Microdurezza Le misure di microdurezza, effettuate presso i laboratori di Alenia Spazio, mostrano una sensibile diminuzione di durezza in corrispondenza del cordone di saldatura, fig. 9. Il valore di durezza Vickers (HV 0.1) del metallo base è pari a circa 155 mentre nella TMAZ e nel nugget si ha una riduzione media di durezza di circa il 35%. L’andamento dei profili di durezza è molto simile sia nel caso di singola che di doppia passata, con una maggiore uniformità dei dati per quanto riguarda la doppia passata. La riduzione di durezza, con il conseguente decadimento delle proprietà meccaniche, è concentrata nel punto terminale della zona termo-meccanicamente alterata, zona in cui si sono tipicamente manifestate sia le rotture statiche che di fatica, con una diminuzione leggermente più accentuata sul lato retreating. È stato dimostrato sperimentalmente, [5], che un aumento della durezza nella zona termicamente alterata, ottenibile agendo sui parametri di saldatura, corrisponde ad un aumento di resistenza meccanica dei giunti.

2.4 Resistenza statica La resistenza statica è stata misurata utilizzando provette di metallo base e provette FSW con saldatura trasversale. Le prove statiche sono state condotte impiegando una macchina servo-idraulica con capacità 25 KN, misurando l’allungamento tramite un estensometro. I risultati medi ottenuti da 3 prove nominalmente identiche hanno condotto ai risultati mostrati in fig. 10. Per il metallo base si è ottenuto un valore medio della tensione di snervamento Sy pari a 549 MPa e una tensione di rottura Su pari a 594 MPa con un allungamento a rottura del 218

5.5%. Per i giunti FSW si è ottenuto per lo snervamento un valore pari a 277 MPa, e 418.25 MPa per la rottura finale, pari rispettivamente al 50 e 70% dei corrispondenti valori del metallo base. Quindi il cosiddetto "coefficiente del giunto" dovrebbe essere valutato sulla base della tensione di snervamento, essendo il parametro maggiormente influenzato dalla saldatura. Per quanto riguarda l’allungamento percentuale, come già descritto nel lavoro precedente [4], si è ottenuto un valore medio dell’allungamento trasversalmente al cordone di saldatura del 10.5%: è importante però osservare che durante le prove statiche si ha un rapido insorgere di deformazioni plastiche all’interno del cordone a causa della minore tensione di snervamento del materiale del cordone di saldatura stesso, mentre il metallo base è ancora in campo elastico. Si possono quindi erroneamente ritenere le fratture dei giunti fragili; in realtà esse sono a tutti gli effetti rotture duttili in cui la plasticizzazione è limitata all’interno del cordone.

2.5 Resistenza a tenso-corrosione Le prove di tenso-corrosione hanno riguardato la lega Al-Li 2195-T8P4, sia metallo base che giunti FSW. In entrambi i casi si è osservata una elevatissima resistenza a tenso-corrosione, come mostrato nella fig. 10, nella quale sono riportati i risultati delle prove statiche condotte su provette esposte ad ambiente corrosivo per 30 giorni e provette esposte allo stesso ambiente corrosivo, contemporaneamente sollecitate con un carico tale da produrre una tensione pari al 75% della tensione di snervamento misurata sulle provette di riferimento (non corrose). L'esposizione all’ambiente corrosivo, anche in presenza di sollecitazione, non riduce la resistenza statica; addirittura per il metallo base si osserva un leggero aumento della tensione di snervamento. Questo fenomeno, "environmetal hardening", osservato anche da altri ricercatori, [6], può essere associato ad una diminuzione della duttilità e della Fracture Toughness. Nel caso in esame il contemporaneo aumento dell’allungamento a rottura sembra escludere questa possibilità. Dopo l'esposizione ad ambiente corrosivo non si sono osservati pits di corrosione nella lega alluminio-litio, per cui si può concludere che in questo materiale anche la resistenza a corrosione generalizzata è molto elevata; prove effettuate in precedenza sul 2219-T851 metallo base e giunzioni VPPA, [7], mostravano una elevata resistenza a tenso-corrosione ma la presenza di pits, particolarmente sulle provette tenso-corrose, era molto comune.

2.6 Resistenza a fatica La resistenza a fatica per il progetto di strutture spaziali viene valutata attraverso le prove e i criteri di meccanica della frattura, come espressamente richiesto dalle normative di progetto. L’effettuazione di alcune prove classiche di fatica è comunque interessante per poter individuare le zone più critiche del giunto, nel quale inserire i difetti iniziali per le prove di meccanica della frattura. In un precedente lavoro [4] si è eseguita una caratterizzazione a fatica dei giunti FSW ottenuti durante la fase di messa a punto del processo. I risultati ottenuti mostravano una elevata resistenza a fatica delle giunzioni che si confrontavano molto bene con i risultati ottenuti per il materiale base e con un comportamento nettamente migliore di altri processi già qualificati in collaborazione con Alenia Spazio come il TIG (Tungsten Inert Gas) o il VPPA (Variable Polarity Plasma Arc), [7]. Nella presente attività i dati relativi al comportamento a fatica sono stati ottenuti analizzando pannelli piani e curvi realizzati con il processo sottoposto a certificazione. I provini tipo dogbone utilizzati per le prove di fatica sono stati ricavati secondo lo schema di fig. 11 da pannelli di spessore 5 mm con saldature longitudinale a singola passata su pannelli piani e pannelli curvi e a doppia passata su pannelli curvi: la sezione di prova di 150 mm2 è centrata su punti del cordone con singola passata, doppia passata o sul punto di chiusura finale, sia per pannelli piani che curvi. Tutte le prove sono state condotte con un martinetto idraulico servocontrollato con capacità di carico di 50 KN. In tutte le prove il rapporto di sollecitazione, R=Smin/Smax, era posto pari a R=0.1. I risultati ottenuti sono riportati in fig. 12. Come si può osservare i dati relativi alle saldature a singola passata presentano una notevole dispersione specialmente per bassi livello di carico; pertanto la linea di best fit è stata tracciata con due curve diverse che sembrano interpretare meglio i risultati. Tale comportamento era stato in effetti rilevato anche per i dati relativi alla prima serie di prove condotte per il processo non ottimizzato, [4]. I provini a doppia passata mostrano invece un comportamento più omogeneo con una diminuzione di prestazioni non significativa sia per quelli ricavati da pannelli piani che dai pannelli curvi, nonostante che nei secondi in alcuni casi si sia osservato un leggero undercut dovuto probabilmente ad una maggiore penetrazione dell’utensile o ad una difficoltà di fissaggio relativo tra i due pannelli curvi da collegare. Pertanto in ogni prova è stato misurato lo spessore effettivo della sezione resistente. I risultati relativi al punto finale di estrazione del pin mostrano una riduzione della resistenza a fatica del giunto in questa posizione. 219

A causa della presenza di alcune discontinuità superficiali descritte in seguito si sono realizzate alcune prove rasando la parte terminale del cordone di saldatura su pannello piano in modo da eliminare il possibile effetto di intaglio superficiale. Le chiusure rasate ricavate da pannelli piani hanno dato risultati confrontabili con le chiusure "as welded" nonostante modalità di frattura differenti. Più eterogenei sono stati i risultati delle chiusure ricavate da pannelli curvi che hanno risentito dei difetti introdotti dall’interazione della spalla dell’utensile con la piastra durante la fase di saldatura, indotti proprio dalla curvatura.

2.7 Analisi delle superfici di frattura Maggiori informazioni si possono trarre dall’analisi delle superfici di frattura osservate nei diversi casi. Come nel lavoro precedente [8] si può osservare, fig. 13a, che le fratture si presentano nella quasi totalità dei casi nella zona termicamente alterata in corrispondenza del punto terminale di contatto della spalla dell’utensile, anche se non si sono riscontrate discontinuità geometriche di rilievo. Le superfici di frattura presentano l’aspetto legnoso caratteristico della lega alluminio-litio già osservato in precedenza nella fase di caratterizzazione del materiale base, [4], poiché la propagazione del difetto avviene fuori della zona meccanicamente alterata. Non si è invece rilevato una debolezza marcata di uno dei due lati del cordone poiché le rotture statisticamente si sono presentate con la stessa probabilità sia sul lato advancing che retreating. In due casi si sono invece osservate fratture nella zona termo-meccanicamente alterata in cui il materiale è stato riomogeneizzato dal processo produttivo, fig. 13b. Comunque in entrambi i casi, prima della rottura finale, i difetti si sono comunque riuniti ad altri piccoli difetti nucleati nella zona termicamente alterata. Per quanto riguarda le chiusure si è osservato in corrispondenza del punto finale di estrazione del pin una traccia, simile ad un grippaggio, le cui dimensioni geometriche fanno supporre che sia prodotta dalla filettatura del perno, fig. 14, da cui nella maggioranza dei casi si sono nucleati i difetti per fatica, fig. 15. Dalla foto emergono chiaramente i punti di innesco delle fratture tipicamente osservate dalle irregolarità visibili anche in superficie. La superficie di frattura, essendo al centro del cordone, mette in evidenza gli anelli concentrici tipici del nugget assimilabili al flusso di materiale che si genera durante la saldatura per l’azione concomitante della rotazione dell’utensile intorno al proprio asse e dell’avanzamento dello stesso lungo la direzione di saldatura. In alcuni provini è stata rasata la superficie del giunto per eliminare l’effetto d’intaglio superficiale. In questo modo si è messo in evidenza un punto debole sul lato rovescio del cordone in cui si sono nucleati i difetti in una fascia di larghezza 2-3 millimetri rispetto all’asse mediano del cordone. Se ne deduce quindi che operativamente la finitura del cordone di saldatura rispetto alla condizione "as welded" non comporti un miglioramento delle prestazioni a fatica. I provini centrati sui punti di chiusura ricavati dai pannelli curvi hanno invece ceduto nella zona di contatto tra spalla e giunto, zona che presentava un leggero undercut, probabilmente causato dalla difficoltà tecnologica di fornire un’opportuna rigidezza della struttura portante del pannello in fase di saldatura.

2.8 Controlli non distruttivi Aspetto fondamentale per la certificazione di un nuovo processo è l’individuazione di tecniche di controllo non distruttivo in grado di rilevare con elevata confidenza gli eventuali difetti presenti all’interno di un giunto. Nel caso del Friction Stir Welding Alenia Spazio intende adottare, unitamente al già collaudato sistema delle "eddy current", un nuovo sistema ad ultrasuoni ad alta definizione, il "phased array". Il controllo con eddy current è stato eseguito tramite una sonda flessibile multi-elemento costituita da 42 bobine che lavorano con sistema multi-plexer, sfruttando gli otto canali dello strumento. La tecnica di ispezione prevede lo spostamento della sonda multi-elemento perpendicolarmente rispetto al cordone di saldatura con l’ausilio di una meccanica manuale fornita di encoders, fig.16. Il software di analisi del sistema eddy currents, alla fine di ogni scansione, fornisce una mappa relativa alla rappresentazione in pianta della saldatura (C-Scan), e la rappresentazione del segnale elettrico. Il metodo "phased array" consiste invece nella scansione elettronica con elementi piezoelettrici che emettono fasci ultrasonori; la sincronizzazione dei fasci ultrasonori emessi dagli elementi viene gestita da un sistema multi-plexer, il quale agisce sia in trasmissione che in ricezione. La tecnica di ispezione prevede infatti l’impiego di una sonda composta da 64 elementi inserita su un adattatore angolato (19°), con un sistema interno di ricircolo dell’acqua movimentata da una meccanica encoderizzata motorizzata biassiale, fig.17. Ogni trasduttore invia un fascio di ultrasuoni ad un determinato angolo ed è focalizzato ad una prefissata profondità. Il sistema software che coordina i vari elementi elabora le informazioni che da essi provengono fornendo in un’unica passata una rappresentazione in pianta della saldatura (C-Scan) e una rappresentazione trasversale del cordone di saldatura (B-Scan). Una parte importante dell’attività di collaborazione con Alenia Spazio è stata finalizzata alla definizione e alla caratterizzazione dei difetti di fatica nei cordoni di saldatura. Per ottenere una banca dati sperimentale di difetti realistici su cui tarare i metodi selezionati, quindici pannelli in 2195-T8P4 saldati sono stati affaticati a diversi livelli. I pannelli utilizzati avevano una larghezza di 200 mm e presentavano una saldatura trasversale, come mostrato in fig.18. Questi pannelli sono stati sottoposti a prova di fatica, Smax=230 MPa, R=0.1; a questa sollecitazione corrisponde una 220 vita media di 109500 cicli, fig. 12. Le prove sono state sospese a numeri cicli compresi fra il 40 e i 60% del numero di cicli a rottura e si sono effettuati i controlli non distruttivi; a tali controlli sono seguiti ulteriori affaticamenti e ulteriori controlli, sia con eddy current che con phased array, in modo da monitorare la crescita dei difetti. In fig. 19 è riportato il numero di cicli a fatica a cui è stato sottoposto ciascun provino indicando schematicamente il tipo di danneggiamento per fatica osservato suddiviso in quattro categorie: nessun danneggiamento riscontrato dai controlli non distruttivi, piccole cricche di fatica (dimensioni inferiori a circa 3 mm), cricche di fatica di dimensioni rilevanti (3-8mm), rottura completa per fatica. I controlli eseguiti hanno evidenziato una maggiore resistenza a fatica dei provini ricavati dalla parte iniziale dei cordoni di saldatura (FXX-1 in fig. 18) e un peggioramento delle prestazioni nei pannelli ricavati dalla parte intermedia e soprattutto dalla parte terminale dei cordone di saldatura (FXX-2 e FXX-3 in fig. 18). Su questi infatti si sono riscontrate anche rotture precoci, fino al 44% della vita stimata (F27-2, F27-3, F28-2 e F28-3, rotti rispettivamente al 50%, 44%, 55% e 60.5% della vita stimata), associate anche ad una più elevata incidenza dei difetti nella TAZ sul lato advancing del cordone, in corrispondenza del punto di contatto della spalla dell’utensile con il giunto saldato. Nei pannelli ricavati dalla parte iniziale delle piastre saldate, FXX-1, non sono stati evidenziati difetti neanche nei pannelli affaticati al 55% della vita. Le successive prove di fatica effettuate non hanno mostrato indizi di possibili variazioni della resistenza lungo i cordoni di saldatura. L’unico dato oggettivo resta l’osservazione di "blisters", piccole aree lucide, assenti nella parte iniziale dei cordoni e progressivamente sempre più presenti lungo le saldature che possono essere indice di un possibile riscaldamento dell’utensile lungo il cordone di saldatura, fig. 8. Le due tecniche di controllo impiegate si sono dimostrate entrambe efficaci e in parte complementari: le eddy current si sono rivelate più adatte in questa applicazione al rilevamento di piccoli difetti in prossimità della superficie esterna, confermando la loro natura di controllo non distruttivo prevalentemente superficiale. La tecnica dei phased array si è invece rivelata ottima per diagnosticare irregolarità all’interno dello spessore del giunto; si è dimostrata invece meno efficace nell’evidenziare piccoli difetti superficiali. Il segnale ultrasonoro risente infatti del disturbo indotto dai "lips" tipici della morfologia superficiale del cordone. La qualità del segnale è però fortemente migliorata levigando il pannello e asportando le irregolarità superficiali che possono coprire le indicazioni provenienti da un difetto. Come esempio di difettologie, la fig. 20a mostra il risultato dell’indagine effettuata con eddy current su un campione contenente un piccolo difetto, avente una dimensione superficiale di 2.5 mm e una profondità di 1 mm. Successivamente il campione è stato rotto staticamente per mostrare il difetto presente. Il difetto è stato chiaramente identificato anche con il metodo dei phased array, come mostrato in fig. 20b: nell’immagine è ben definita la cricca di fatica ma la tecnica necessita ancora di una messa a punto definitiva per evitare il disturbo visibile sui lati del cordone.

3. CONCLUSIONI

L’ampia attività di ricerca condotta ha evidenziato le notevoli potenzialità del processo di saldatura solida Friction Stir Welding in associazione con la nuova lega alluminio-litio WeldaliteTM. La giunzione analizzata non presenta difettologie al suo interno come confermato dalle ottime proprietà meccaniche, analizzate anche in ambiente aggressivo, che si confrontano positivamente con i processi di saldatura già impiegati da Alenia Spazio. La sollecitazione a fatica non ha evidenziato zone particolarmente deboli nel giunto, compresa anche la tipica zona critica della chiusura finale. I difetti nucleatisi non sono correlati a sistematiche carenze del processo di saldatura e sono comunque stati chiaramente individuati con i sistemi NDI selezionati. Pertanto il processo risulta estremamente efficace per la realizzazione di pressur vessels di impiego spaziale da un punto di vista strutturale ed economicamente vantaggioso dati i ridotti tempi di realizzazione associati. L’attività si è conclusa con la costruzione del serbatoio dimostrativo per ossigeno liquido FAST 2, fig. 21. In conseguenza dei risultati estremamente promettenti è allo studio presso Alenia Spazio un’estensione della tecnologia Friction anche per applicazioni sulla lega 2219. In particolare la saldatura Friction è stata introdotta nel progetto preliminare di realizzazione del CEV (Cabin Exploration Vehicle), il veicolo spaziale pressurizzato per il trasporto umano sulla stazione spaziale con possibile estensione alle missioni abitate sulla Luna e Marte.

221

BIBLIOGRAFIA

[1] W. M. Thomas, E. D. Nicholas, J. C. Needham, M. G. Church, P. Templesmith, C. J. Dawes International, Patent Application N°. PCT/IGB92/02203 and GB Patent Application No. 9125978.9, 1991. [2] M. Nebiolo, L. Basile, A. Cramarossa, R. Laudato, A. Suppo, M. F. Rossi "Cryogenic Tanks for Launch Vehicles" 55th International Astronautical Congress of the International Astronautical Federation, the International Academy of Astronautics, and the International Institute of Space Law, Vancouver, Canada, Oct. 4-8, 2004. Paper IAC-04-S.3.10. [3] R. J. Ding, U.S.Patent No. #5,893,507 U.S. Patent, NASA Marshall Space Flight Center, USA, 2001. [4] A. Lanciotti, C.Polese "Alcune proprietà di giunzioni 'Friction Stir Welding' in due leghe di alluminio di impiego aerospaziale" Memoria presentata al Convegno "Friction Stir Welding per Applicazioni Aerospaziali"; Pisa 1 Aprile 2003. [5] G. Biallas, R. Braun, C. Dalle Donne, G. Staniek, and W. A. Kaysser, Mechanical Properties and Corrosion Behavior of Friction Stir Welded 2024-T3, 1st International Symposium on Friction Stir Welding, TWI, UK, 1999. [6] W. Hu, E. I. Meletis "Corrosion and environment-assisted cracking behaviour of friction stir welded Al 2195 and Al 2219 alloys". Material Science Forum, Vol. 331-337, pp. 1683-1688, 2000. [7] A. Lanciotti, A. Belmondo "Mechanical properties of Al 2219 VPPA weldments". 48th International Astronautical Congress, Turin, 1997. Paper IAF-97-1.4.07. [8] A. Lanciotti, C. Polese "Resistenza a Fatica e Caratterizzazione Damage Tolerance di Giunzioni Friction Stir Welding nella Lega Alluminio-Litio 2195-T8", Giornata di Studio "Friction Stir Welding Day" Torino, 15 Giugno 2004.

Ringraziamenti

Gli autori ringraziano per l’attiva collaborazione il Sig. Angelo Camassa, il Sig. Valter Capitani e la Dott.ssa Monica Bosco del Laboratorio Materiali e Processi di Alenia Spazio e l’Ing. Federico Ferrari per l'aiuto fornito nell’esecuzione delle prove sperimentali e l’elaborazione dei risultati ottenuti. Per lo svolgimento di questa attività Alenia Spazio ha usufruito di un finanziamento dell’ASI, Agenzia Spaziale Italiana, per il programma FAST 2 e il Dipartimento di Ingegneria Aerospaziale ha usufruito di un finanziamento del Ministero dell’Istruzione, dell’Università e della Ricerca, programma COFIN 2003/2004.

222

Fig. 1 – Il progetto del pressure vessel Alenia Spazio Fast 2

5 m 2.2 1. 3 m

Fig. 2 – L’utensile Friction Stir Welding

Fig. 3 – Il “retractable Pin Tool”

a b

Fig. 4 –Punto di chiusura: a) “retractable pin tool” e b) “key-hole” classico 223

Fig. 5 – Dinamica di estrazione del “retractable pin tool”

Fig. 6 – Particolare di una saldatura FSW circonferenziale

Singola passata

TMAZ NUGGET MB 5 mm HAZ a

TMAZ NUGGET 5 mm MB HAZ b Doppia passata Fig. 7 – Macro-micrografia di un cordone FSW: a) singola passata, b) doppia passata 224

1 mm

Fig. 8 – Presenza di “blisters” nella zona centrale e terminale dei cordoni di saldatura

Fig. 9 – Profili di microdurezza 225

Fig. 10 – Proprietà meccaniche generali, metallo base e giunti FSW

Fig. 11 – Schema realizzativo provini FSW 226

R=0.1 300 Materiale base

250 Passata doppia

[MPa] Passata singola max 200 S

Passata singola Passata doppia Chiusura Chiusura Materiale base

150 104 105 106 Cycles Fig. 12 – Risultati delle prove di fatica effettuate sui giunti FSW e sul materiale base

1 mm

0.5 mm

Fig. 13 – Tipiche rotture per fatica 227

Fig. 14 – Dettaglio del punto di chiusura di un giunto saldato

Fig. 15 – Dettaglio di una rottura per fatica in un punto di chiusura 228

Fig. 16 – Strumentazione eddy current

Fig. 17 – Strumentazione phased array

Fxx-1 Inizio saldatura Fxx-2 Fxx-3 300

200 850

Fig. 18 – Prove di fatica. Identificazione dei provini utilizzati per la calibrazione del sistema di controllo dei giunti saldati 100000 229 Smax=230 MPa; R=0.1 Assenza di danneggiamento 80000 X 2c=1-3 mm XX 2c=3-6 mm X XXX Rottura completa 60000 XXX XXX X XX XX XX XX Cicli XX XXX 40000 XX

20000

0 123 Pannello n.

Fig. 19 – Sintesi dei risultati delle prove di fatica effettuate sui pannelli disturbo nel segnale b)

C-scan

B-scan a)

Fig. 20 – Individuazione di un difetto con a) eddy current, b) phased array 230

Fig. 21 – Il serbatoio criogenico dimostrativo di Alenia Spazio Fast 2 PUBLICATIONS I

5. PUBLICATIONS

CONVENTIONAL STRUCTURES

SECONDARY BENDING

[1] A. Lanciotti, C. Polese “Fatigue crack propagation of through cracks in thin sheets under combined traction and bending stresses” Fatigue and Fracture of Engineering Materials and Structures , v. 26, 2003, pp. 421-428.

Valutazione del prodotto Codice: 039816 Panel: 09 - Ingegneria industriale e dell'informazione Categoria ISI: IS0305 - Aerospace Engineering Giudizio finale: Eccellente

[2] A. Lanciotti, C. Polese “Valutazione numerica della propagazione di difetti sotto l'azione combinata di carichi di trazione e flessione” XVII Congresso della Associazione Italiana di Aeronautica e Astronautica, Roma, 15-19 Settembre 2003.

TECHNIQUES TO ENHANCE FATIGUE LIFE OF STRESSED HOLES

[3] Competitive and Sustainable Growth Programme, Advanced Design concepts and Maintenance by Integrated Risk Evaluation for aerostructure, acronym ADMIRE, 5th Framework Programme, Project N°GRD1-2000-25069, Contratto N°G4RD-CT-2000- 00396:

o G. Cavallini, R. Lazzeri, C. Polese, “WP3 Activity at University of Pisa”, Technical Report, ADMIRE Atene 20-21 June 2002. o G. Cavallini, R. Lazzeri, C. Polese, “UP Contribution to WP3.1.1 ADMIRE activity”, Technical Report, 15 May 2003. o G. Cavallini, R. Lazzeri, C. Polese “Final Report on Tests Carried out for the Evaluation of Short Crack Effect”. Document N°. ADMIRE-TR-3.0-61-3.1/UP., New Perspectives in Aeronautics, 31 July 2003.

[4] A. Lanciotti, C. Polese “The effect of interference fit fasteners on the fatigue life of central hole specimens” Fatigue and Fracture of Engineering Materials and Structures, vol. 28, n. 8, 2005, pp. 587-597.

231 PUBLICATIONS I

[5] A. Lanciotti, C. Polese “Effetto dell'interferenza rivetto-foro sulla resistenza a fatica di provini con e senza trasferimento di carico" XVIII Congresso della Associazione Italiana di Aeronautica e Astronautica, Volterra, 19-22 Settembre 2005. [6] L. Boni, A. Lanciotti, R. Gonella, F. Pietra, C. Polese “Numerical simulations and experimental evaluations of some technological processes employed to improve the fatigue resistance of stressed holes” TCN-CAE 2005, International Conference on CAE and Computational Technologies for Industry, 5-8 October 2005, Lecce, Italy.

[7] A. Lanciotti, L. Boni, C. Polese “Residual stress field in open hole specimens subjected to “Cold Working” and its effect on the fatigue resistance” Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2006/4, Febbraio 2006.

[8] A. Lanciotti, L.Boni, C. Polese “Fatigue Resistance of Riveted Joints with Cold Worked Holes”. Collaborazione di ricerca fra il Dipartimento di Ingegneria Aerospaziale e la Augusta Westland Elicotteri, final report, Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2006-14, June 2006.

[9] F. Franchini, F. Pietra, C. Polese, S. Urso Sculco “Chaboche Material Model Optimization for Evaluation of Residual Stress Field in Cold Worked Holes” International ModeFRONTIER Users’ Meeting, 28-29 September 2006, Trieste, Italy.

[10] L.Boni, A. Lanciotti, C. Polese “Numerical and experimental assessment of fatigue behaviour of “cold worked” open holes”, Invited Lecture at the 17° Italian ABAQUS Users’ Meeting, October 04-06 2006, Pisa – Italy

[11] F. Franchini, F. Pietra, C. Polese, S. Urso Sculco “Finite element analysis and experimental evaluation of fatigue life enhance of Cold Worked holes” ANSYS Italian Users’ Meeting, 9-10 November, Stezzano (BG), Italy.

[12] L.Boni, A. Lanciotti, C. Polese “Fatigue and Crack propagation in Open Hole Specimens with Cold Worked Holes”, 24th ICAF Symposium, 16-18 May 2007 Naple, Italy (in progress)

[13] A. Lanciotti, F.Nigro, C.Polese “Caratterizzazione Damage Tolerance dell'attacco ala- fusoliera del velivolo M346” XVII Congresso della Associazione Italiana di Aeronautica e Astronautica, Roma, 15-19 Settembre 2003.

232 PUBLICATIONS I

[14] A. Lanciotti, C. Polese, L. Cabiddu “Attività sperimentale di caratterizzazione Damage Tolerance dell'attacco ala-fusoliera del velivolo M346”. Contratto di ricerca fra il Dipartimento di Ingegneria Aerospaziale ed AerMacchi S.p.A., Venegono Superiore. Relazione finale. Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2004/9, Giugno 2004.

[15] A. Lanciotti, F. Nigro, C. Polese “Damage Tolerance evaluation of the wing to fuselage connection of the new trainer aircraft M346”. Fatigue and Fracture of Engineering Materials and Structures, October 2006.

INTEGRAL STRUCTURES

[16] A. Lanciotti, C. Polese “Alcune proprietà di giunzioni “Friction Stir Welding” in due leghe di alluminio di impiego aerospaziale” Memoria presentata al Convegno “Friction Stir Welding per Applicazioni Aerospaziali”, Pisa, 1 Aprile 2003.

[17] Lanciotti, C. Polese “Resistenza a fatica di elementi in acciao inox 301 saldati per punti” Contratto di ricerca fra il Dipartimento di Ingegneria Aerospaziale e la Breda Costruzioni Ferroviarie. Relazione finale. Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2004/8, Maggio 2004.

[18] A. Lanciotti, C. Polese “Resistenza a fatica e caratteristiche di Damage Tolerance di giunzioni Friction Stir Welding nella lega alluminio-litio 2195-T8” Giornata di Studio: “Friction Stir Welding Day”, Torino, 15 Giugno 2004.

[19] A. Lanciotti, C. Polese “Impact of Direct Water Cooling on the quality and performance of Friction Stir Welded joints in an AlMgSiCu aluminium alloy” Atti del Dipartimento di Ingegneria Aerospaziale, Università di Pisa, ADIA 2004-2, Settembre 2004.

[20] A. Lanciotti, C. Polese “Alcuni aspetti della resistenza a fatica e delle caratteristiche di Damage Tolerance di giunzioni Friction Stir Welding in una lega alluminio-litio” Collaborazione di ricerca fra il Dipartimento di Ingegneria Aerospaziale ed Alenia Spazio. Relazione n. 1. Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2004/1, Ottobre 2004.

[21] A. Lanciotti, R. Laudato, C. Polese “Proprietà meccaniche di giunzioni “Friction Stir Welding” in una lega alluminio-litio per applicazioni spaziali”. XVIII Conferenza della Associazione Italiana di Aeronautica e Astronautica, Volterra, 19-22 Settembre 2005.

[22] T. Ghidini, C. Polese, C. Dalle Donne, A. Lanciotti “Ermüdungsrißwachstum bei Reibrührschweißnähten unter Flugbedingungen”. DVM-Bericht 132, pp. 119-129, 2005.

233 PUBLICATIONS I

[23] A. Lanciotti, C. Polese “Fracture Tests on 2219-T851 VPPA welded panels” Collaborazione di ricerca fra il Dipartimento di Ingegneria Aerospaziale ed Alcatel Alenia Space Italia. Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2005/32, Dicembre 2005. [24] A. Lanciotti, C. Polese “Damage Tolerance properties in 2219-T851 Friction Stir Welded panels” Collaborazione di ricerca fra il Dipartimento di Ingegneria Aerospaziale ed Alcatel Alenia Space Italia. Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2005/33, Dicembre 2005.

[25] A. Lanciotti, C. Polese “Resistenza a frattura ed a fatica di giunzioni Friction Stir Welding in una lega alluminio-litio: alcuni risultati integrativi”. Collaborazione di ricerca fra il Dipartimento di Ingegneria Aerospaziale ed Alcatel Alenia Space Italia. Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2006/1, Gennaio 2006.

[26] T. Ghidini, C. Polese, C. Dalle Donne, A. Lanciotti “Simulations of fatigue crack propagation in Friction Stir Welds under flight loading conditions”, Materials Testing 7/8, pp. 370-375, 2006.

[27] Innovative Fatigue and Damage Tolerance Methods for the Application of New Structural Concepts, acronym DaToN, 6th Framework Programme, Priority 4, Aeronautics and Space, Contract no. FP6-516053, related contract no. AST3-CT-2004-516053:

o Lanciotti, L. Lazzeri, C.Polese “ Working Document WP1” Braunschweig 18-19 April 2005.

o Lanciotti, L. Lazzeri, C.Polese “Proposal for a WP1 Input” DaToN WP 1.1 – 1.0, Pisa 27-28 October 2005.

o Lanciotti, L. Lazzeri, C.Polese “Preliminary calculations of crack propagation in integrally stiffened panels” DaToN WP 2.1 – 1.0, Paris, April 2006.

o Lanciotti, L. Lazzeri, C.Polese “Preliminary calculations of crack propagation in integrally stiffened panels” DaToN WP 2.1 – 2.0, Emmeloord, September 2006.

FIBRE METAL LAMINATES

[28] A. Lanciotti, C. Polese “Crack propagation and fracture toughness tests of fibre metal laminates, (Glare)” Contratto di ricerca fra il Dipartimento di Ingegneria Aerospaziale ed Alenia Aeronautica. Relazione finale. Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2003/7, Maggio 2003.

234 PUBLICATIONS I

[29] A. Lanciotti, C. Polese “Mechanical properties of fibre metal laminates tested at different temperatures” Contratto di ricerca fra il Dipartimento di Ingegneria Aerospaziale ed Alenia Aeronautica. Relazione finale. Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2003/18, Novembre 2003.

ADDITIONAL ACTIVITIES

[30] A. Lanciotti, C. Polese “Static test on the nose landing gear drag brace lower arm of the ATR 42/72 aircraft”. Contratto di ricerca fra il Dipartimento di Ingegneria Aerospaziale e la Società ATR, Toulose. Relazione finale. Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2004/4, Febbraio 2004. [31] A. Lanciotti, C. Polese “Indagine sulla rottura per fatica di alcuni carrelli ferroviari” Contratto di ricerca fra il Dipartimento di Ingegneria Aerospaziale e la Breda Costruzioni Ferroviarie. Relazione finale. Documenti del Dipartimento di Ingegneria Aerospaziale DDIA 2004/7, Aprile 2004.

INTERNATIONAL CONFERENCES

[32] A. Salvetti and L. Lazzeri “Review of aeronautical fatigue investigations carried out in Italy during the period April 2001 - March 2003” chapter 1-2-3-4-6-8, A.I.F.A. - ITALIAN ASSOCIATION FOR FATIGUE IN AERONAUTICS, Department of Aerospace Engineering University of Pisa - Italy , 28th ICAF Conference, Lucerne, Switzerland, May 5-6 2003.

[33] A. Salvetti, L. Lazzeri, Department of Aerospace Engineering University of Pisa - Italy “Review of aeronautical fatigue investigations carried out in Italy during the period April 2003 - March 2005”, 29th ICAF Conference, Hamburg, Germany, June 6-7 2005.

[34] Claudio Dalle Donne, EADS Corporate Research Center Germany, Metallic Structures, “Review of Aeronautical Fatigue Investigations in Germany During the Period March 2003 to May 2005”, 29th ICAF Conference, Hamburg, Germany, June 6-7 2005.

235 TUTOR ACTIVITY: FINAL THESIS IN AEROSPACE ENGINEERING I 6. TUTOR ACTIVITY: FINAL THESIS IN AEROSPACE ENGINEERING

OCTOBER 2002

1) Gregorio Nazzani, Emanuele Vincentini "ANALISI STRUTTURALE DELL'ALBERATURA DI UNA IMBARCAZIONE DI GRANDI DIMENSIONI"

FEBRUARY 2003

2) Francesco Pietra "EFFETTI DELL’INTERFERENZA E DELLA FLESSIONE SECONDARIA SULLA RESISTENZA A FATICA DI GIUNZIONI RIVETTATE"

3) Federico Tarabella (Rif. Bibl. DIA: 418) "VALUTAZIONE NUMERICA DELLA PROPAGAZIONE DI CRICCHE SOTTO CARICHI COMBINATI DI TRAZIONE E FLESSIONE"

APRIL 2003

4) Guido Becchi (Rif. Bibl. DIA: 459) "RESISTENZA A FATICA DI GIUNZIONI “FRICTION STIR WELDING” IN UNA LEGA ALLUMINIO-LITIO DI NUOVA GENERAZIONE"

5) Elisabetta Fattorini (Rif. Bibl. DIA: 391-395) "CARATTERIZZAZIONE DI GIUNZIONI IN LEGA 6082-T6 SALDATE FRICTION STIR WELDING"

JULY 2003

6) Fabio Forlani, Gianantonio Toniolo (Rif. Bibl. DIA: 457) "CARATTERIZZAZIONE DAMAGE-TOLERANCE DELL’ATTACCO ALA- FUSOLIERA DI UN VELIVOLO DI ADDESTRAMENTO"

7) Roberto Marcalli "PROPAGAZIONE DI DIFETTI E RESISTENZA STATICA RESIDUA NEI LAMINATI METALLICI RINFORZATI CON FIBRA DI VETRO"

OCTOBER 2003

8) Igor Azara (Rif. Bibl. DIA: 444) "STUDIO DEL COLLEGAMENTO ALA-FUSOLIERA SMONTABILE PER IL VELIVOLO PIAGGIO P180 AVANTI"

236 TUTOR ACTIVITY: FINAL THESIS IN AEROSPACE ENGINEERING I 9) Giuseppe Caruso (Rif. Bibl. DIA: 426-455) "PROVE DI TENUTA SU FUEL RAIL DI BASSA PRESSIONE PER AUTOVETTURE A BENZINA"

10) Vittorio Cipolla, PierPaolo Pergola "ALCUNE PROPRIETÀ MECCANICHE DI GIUNZIONI SALDATE FRICTION STIR IN LEGA AL 6013-T6"

11) Damiano Barletta "GIUNZIONI INCOLLATE DI IMPIEGO AERONAUTICO"

DECEMBER 2003

12) Daniele Fois "VALUTAZIONE DEL SOFTWARE CATIA V5 PER PIAGGIO AERO INDUSTRIES"

FEBRUARY 2004

13) Daniele Caprai "ALCUNE PROPRIETA’ MECCANICHE DEI LAMINATI METALLICI RINFORZATI CON FIBRE (F.M.l.) "

14) Luca Romeo "FRICTION STIR WELDING"”

APRIL 2004

15) Lorenzo Cabiddu "CARATTERIZZAZIONE “DAMAGE TOLERANCE” TRAMITE ANALISI NUMERICO-SPERIMENTALE DELL’ATTACCO ALA FUSOLIERA DI UN ADDESTRATORE DI VOLO"

16) Nicola Pasquale "EFFETTI DELL’INTERFERENZA SULLA RESISTENZA A FATICA IN UN ACCOPPIAMENTO PERNO-FORO"

JUNE 2004

17) Sara Emanuela Giovannetti, Francescostefano Mazzeo "ALCUNI ASPETTI DELLA RESISTENZA A FATICA DI GIUNZIONI SALDATE PER PUNTI NELL’ACCIAIO AISI 301"

18) Daniele De Vita "LA PALLINATURA"

237 TUTOR ACTIVITY: FINAL THESIS IN AEROSPACE ENGINEERING I 19) Marco DeTata "TECNICHE DI SALDATURA E TAGLIO LASER"

JULY 2004

20) Stefano Androni "METODI NON DISTRUTTIVI PER LA RIVELAZIONE DELLA CORROSIONE NELLE STRUTTURE AERONAUTICHE"

21) Carmela Fierro "FRICTION STIR WELDING"

OCTOBER 2004

22) Francesco Fruci "EFFETTI DELL’INTERFERENZA SULLA RESISTENZA A FATICA DELLE GIUNZIONI RIVETTATE"

NOVEMBER 2004

23) Luca Agostini, Chiara Battistelli (Rif. Bibl. DIA: 490) "APPLICAZIONE DEI MATERIALI COMPOSITI IN AMBITO NAVALE”

24) Paola Apruzzese (Rif. Bibl. DIA: 506) "AUTOTENSIONI CONSEGUENTI A COLD WORKING E INTERFERENZA, EFFETTI SULLA RESISTENZA A FATICA"

25) Federico Ferrari "PROPRIETÀ DI DAMAGE TOLERANCE DI GIUNZIONI “FRICTION STIR WELDING” IN UNA LEGA ALLUMINIO LITIO PER IMPIEGO SPAZIALE"

JULY 2005

26) Gianluca Magagnini "ANALISI DELLE CARATTERISTICHE DI RESISTENZA STATICA E DI FATICA DI ALCUNI TIPI DI GIUNTI IN GLARE"

27) Biancamaria Felice "CARATTERIZZAZIONE “DAMAGE TOLERANCE” DELL’ATTACCO ALA- FUSOLIERA IN LEGA D’ALLUMINIO 7050 -T7451 DI UN ADDESTRATORE DI VOLO TRAMITE ANALISI NUMERICO-SPERIMENTALE"

238 TUTOR ACTIVITY: FINAL THESIS IN AEROSPACE ENGINEERING I OCTOBER 2005

28) Alessio Iantorno "RESISTENZA A FATICA DI GIUNZIONI SALDATE PER PUNTI IN UN ACCIAIO INOSSIDABILE. RISULTATI SPERIMENTALI E POSSIBILI ESTENSIONI A GEOMETRIE DIVERSE"

DECEMBER 2005

29) Giuseppe Allegra, Emanuele Calvagna "NUCLEAZIONE E PROPAGAZIONE DEI DIFETTI NEI FORI TRATTATI CON IL COLD WORKING"

30) Saverio Picerno (Rif. Bibl. DIA: 534) "ANALISI DEI MODI DI ROTTURA PER FATICA IN FORI TRATTATI CON COLD WORKING"

APRIL 2006

31) Lia Ruggeri "NUCLEAZIONE E PROPAGAZIONE DI DIFETTI IN FORI TRATTATI CON “STRESS WAVE” "

JUNE 2006

32) Contini Francesco, Domenico Zaffino "RESISTENZA A FATICA E RESISTENZA STATICA RESIDUA SU PROVINI FML LAP JOINT"

JULY 2006

33) Ivan Andreani, Vjola Ristori "PROPAGAZIONE DI DIFETTI IN GIUNZIONI IN LEGA DI ALLUMINIO 2219 SALDATI “FRICTION STIR”"

34) AndreaSebastian Daquino, Manuel D’Eusebio "MIGLIORAMENTO DELLA RESISTENZA A FATICA E ANALISI DELLE TENSIONI RESIDUE SU FORI TRATTATI CON LA TECNICA DEL “COLD WORKING”"

35) Marco Dezi "RESISTENZA A FRATTURA E PROPAGAZIONE DI DIFETTI IN GIUNTI SALDATI CON TECNICHE FRICTION STIR E VARIABLE POLARITY PLASMA ARC NELLA LEGA 2219-T851"

239 TUTOR ACTIVITY: FINAL THESIS IN AEROSPACE ENGINEERING I 36) Francesco Contini "RESISTENZA A FATICA E RESISTENZA STATICA RESIDUA SU PROVINI FML LAP JOINT"

OCTOBER 2006

37) Gabriella Perrone "ANALISI MEMBRANALE DI UNA VELA SOLARE"

38) Filippo Nogarin "STUDIO DI SISTEMA E PROGETTAZIONE DI UN PICCOLO AEROGENERATORE (15 KW)"

NOVEMBER 2006

39) Nicola Pasquale "PROPAGAZIONE DI DIFETTI IN PANNELLI IRRIGIDITI"

40) Daniele Dignani, Stefano Maestrini "ANALISI DELLE CARATTERISTICHE DI RESISTENZA STATICA ED A FATICA DI GIUNZIONI IN GLARE A DUE FILE DI RIVETTI"

41) Teresa Spadafora "STRESS CORROSION"

IN PROGRESS FINAL THESIS

42) Sergio Urso Sculco "ANALISI FEM DEL PROCESSO DI COLD WORKING CON MODELLO CHABOCHE"

43) Anita Catapano, Sandro Sansica "SPLIT MANDREL COLD WORKING"

44) Luca Niccolini "VALUTAZIONI NUMERICHE E SPERIMENTALI DI LAMINATI GLARE"

240