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Flame straightening is a manufacturing process for achieving geometry of elements and members made up of structural steel. The technical background on flame straightening in the workshops of steel constructors is often based only on empirical knowledge, if it exists at all. As a consequence, the straightening of steel construction elements to achieve the required geometrical shape absorbs a large part of the manufacturing costs. This is due to uncertainty about the correct flame straightening process and the lack of background knowledge on its effects. Even though in scientific circles there is some knowledge on the mechanisms of flame straightening for different temperatures, holding times and steel grades, the knowledge is scattered, not well documented and has not been transferred to complex steel structures, sections, stiffness and real (extended) geometries. Also, application techniques, flame straightening procedures and an insight parameter clarification in particular for high strength steels do not exist at all.

This report presents the results of the European research project Optistraight. Through experimental, numerical and analytical investigations, the mechanisms of different flame straightening processes have been clarified. Together with the available, but scattered, knowledge on this fabrication process the results give an in-depth view of the flame straightening process. Based upon this knowledge, prediction means have been developed to estimate the straightening result beforehand and to avoid expensive ‘trial-and-error’ tests, detrimental impacts on the material or excessive energy inputs.

Optimisation and improvement of the flame straightening process

(Optistraight)

doi:2777/37733

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25120

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European Commission

Research Fund for Coal and SteelOptimisation and improvement of the flame straightening process

(Optistraight)

D. SchäferRWTH Aachen University, Institute for Steel StructuresMies-van-der-Rohe-Straße 1, 52074 Aachen, GERMANY

V. RinaldiArcelorMittal Research and Development

Rue de Luxembourg, 66, 4009 Esch-sur-Alzette, LUXEMBOURG

D. Beg, P. MožeUniversity of Ljubljana, Faculty of Civil and Geodetic Engineering, Chair for Metal Structures

Jamova cesta 2, SI-1000 Ljubljana, SLOVENIA

R. Lacalle, J. Portilla, D. Ferreño, J. A. ÁlvarezUniversity of Cantabria, Departamento de Ciencia e Ingeniería del Terreno y de los Materiales,

E.T.S. Ingenieros de CaminosAvenida de los Castros s/n, 39005 Santander, SPAIN

R. Willms, J. SchützAG der Dillinger Hüttenwerke

Werksstraße 1, 66763 Dillingen/Saar, GERMANY

Contract No RFSR-CT-2007-00040 1 July 2007 to 30 June 2010

Final report

Directorate-General for Research and Innovation

2012 EUR 25120 EN

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Luxembourg: Publications Office of the European Union, 2012

ISBN 978-92-79-22426-3

doi:10.2777/37733

ISSN 1831-9424

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Table of Contents

Table of Contents....................................................................................................................................3 

Summary .................................................................................................................................................7 

Scientific and technical description of the results..............................................................................13 

Comparison of initially planned activities and work accomplished .......................................................13 

1  Introduction ......................................................................................................................................15 

1.1  Overview of the project ................................................................................................................15 

1.2  Objectives of the project ..............................................................................................................17 

2  Experimental investigations ............................................................................................................19 

2.1  Tests with small scale specimen ..................................................................................................19 

2.1.1  Tests on partly heated specimen .............................................................................................. 19 

2.1.2  Free ends completely heated .................................................................................................... 29 

2.1.3  Material investigation on small scale specimen ....................................................................... 30 

2.1.4  Residual stress measurement on small scale specimen ............................................................ 37 

2.2  Large scale tests ...........................................................................................................................39 

2.2.1  Introduction .............................................................................................................................. 39 

2.2.2  Testing of long profiles ............................................................................................................ 40 

2.2.3  Testing of plate material .......................................................................................................... 44 

2.2.4  Tests on plates with longitudinal and transversal stiffeners ..................................................... 45 

2.2.5  Testing of distortions of attachment, profiles and singles plates ............................................. 48 

2.2.6  Tests on ends of profiles .......................................................................................................... 53 

2.2.7  Material investigation on large scale tests ............................................................................... 55 

2.2.8  Residual stress measurement on beams ................................................................................... 58 

3  Numerical investigations .................................................................................................................61 

3.1  Numerical model of the flame ......................................................................................................61 

3.1.1  Heat input and heat flow density distribution parameters in the literature .............................. 62 

3.1.2  Experimentally derived heat input ........................................................................................... 63 

3.2  Plate structures .............................................................................................................................63 

3.2.1  The selection of an efficient finite element .............................................................................. 63 

3.2.2  Numerical simulations of plate elements ................................................................................. 64 

3.2.3  Numerical simulations of T elements ...................................................................................... 67 

3.3  Bar shaped structures ...................................................................................................................71 

3.3.1  Transient, fully coupled thermo-mechanical numerical simulations of beams ........................ 71 

3.3.2  Steady-state, fully coupled thermo-mechanical numerical simulations of beams ................... 75 

3.3.3  Methodology for 3D-modeling of the flame bending process of beams .................................. 76 

4  Analytical investigations ..................................................................................................................81 

4.1  Presentation of analytical model for bar shaped structures ..........................................................81 

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4.1.1  Model 1: Critical temperature approach .................................................................................. 81 

4.1.2  Model 2: Determination of the deformation of bar-shaped elements by flame straightening using analytical prediction means ...................................................................................................... 83 

4.2  Comparison of analytical and numerical prediction means .........................................................89 

5  Conclusions .......................................................................................................................................91 

5.1  Conclusions of experimental investigations .................................................................................91 

5.2  Conclusions of numerical investigations......................................................................................93 

5.3  Conclusions of analytical investigations ......................................................................................94 

5.4  General conclusion .......................................................................................................................94 

6  Guideline ...........................................................................................................................................97 

7  Exploitation and impact of research results ..................................................................................99 

8  List of references ............................................................................................................................101 

9  List of figures ..................................................................................................................................105 

10 List of tables ....................................................................................................................................111 

11 List of acronyms and abbreviations .............................................................................................113 

Appendix 1: List of documents distributed in the frame of OPTISTRAIGHT ............................115 

Appendix 2: Guideline .......................................................................................................................117 

1  Explanation of mechanism ............................................................................................................117 

2  Equipment, adjustment of energy and oxygen partition of the torch .......................................118 

2.1  Heat input ...................................................................................................................................118 

2.1.1  Gas ......................................................................................................................................... 118 

2.1.1.1  Acetylene ......................................................................................................................... 118 

2.1.1.2  Propane ............................................................................................................................ 118 

2.1.2  Torch typology ....................................................................................................................... 118 

2.1.2.1  Single orifice torch tip ..................................................................................................... 119 

2.1.2.2  Multiple orifice torch tip .................................................................................................. 119 

2.2  Measurement equipment ............................................................................................................119 

2.2.1  Temperature control ............................................................................................................... 119 

2.2.1.1  Thermocouples ................................................................................................................ 119 

2.2.1.2  Contact thermometers ...................................................................................................... 119 

2.2.1.3  Optical and infrared pyrometers ...................................................................................... 120 

2.2.1.4  Thermographic cameras................................................................................................... 120 

2.2.1.5  Temperature crayons ....................................................................................................... 120 

2.2.1.6  Visual control .................................................................................................................. 120 

2.2.2  Displacement control ............................................................................................................. 120 

2.2.2.1  Simple measuring tool ..................................................................................................... 120 

2.2.2.2  Inductive displacement transducer .................................................................................. 120 

2.2.2.3  Laser displacement transducer ......................................................................................... 121 

2.3  Additional loads or restraints .....................................................................................................121 

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2.3.1  Mechanical jack ..................................................................................................................... 122 

2.3.2  Hydraulic jack ........................................................................................................................ 122 

2.3.3  Dead loads .............................................................................................................................. 122 

2.3.4  Location of loads or restraints ................................................................................................ 122 

3  Heating patterns .............................................................................................................................123 

3.1  Types of heating patterns ...........................................................................................................123 

3.1.1  Heat spots ...................................................................................................................................123 

3.1.2  Heat stripe, line heat ...................................................................................................................124 

3.1.3  Banded heat ................................................................................................................................125 

3.1.4  Vee heat ......................................................................................................................................126 

4  Procedure ........................................................................................................................................128 

5  Quantification and control of heat input .....................................................................................129 

5.1  Quantification of effective heat input.........................................................................................129 

5.2  Testing procedure to quantify the effective heat input ...............................................................129 

5.3  Calibrating the heat input by numerical means ..........................................................................130 

6  Avoiding an impact on the material .............................................................................................133 

6.1  Temperature range for flame straightening ................................................................................133 

6.2  Delivery conditions ....................................................................................................................133 

6.2.1  As rolled (AR)........................................................................................................................ 133 

6.2.2  Normalised (N) ...................................................................................................................... 133 

6.2.3  Quenched and tempered (Q) .................................................................................................. 133 

6.2.4  Thermomechanically rolled (M) ............................................................................................ 134 

6.2.5  Microstructure ........................................................................................................................ 134 

6.3  Upper limit of applicable temperature .......................................................................................135 

7  Prediction means and examples ....................................................................................................136 

7.1  Application of prediction means for plates with line heats ........................................................136 

7.1.1  Procedure for the prediction of the rotation angle ................................................................. 136 

7.1.2  Example ................................................................................................................................. 138 

7.2  Application of analytical model 1 to I and H profiles ................................................................140 

7.2.1  Strong axis bands ................................................................................................................... 141 

7.2.2  Weak axis bands – V heat ...................................................................................................... 143 

7.3  Application of numerically derived prediction means for beams ..............................................143 

7.3.1  Influence of temperature ........................................................................................................ 143 

7.3.2  Influence of bending moment ................................................................................................ 144 

7.3.3  Example 1: Vee heat about the strong axis ............................................................................ 146 

7.3.4  Example 1: Vee heat about the weak axis .............................................................................. 147 

7.4  Further prediction means for beams (model 2) ..........................................................................148 

8  Quantification of displacement .....................................................................................................149 

Appendix 3: Guideline for the numerical simulation of flame straightening ...............................151 

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1  Material model ....................................................................................................................... 151 

2  Numerical model of flame ..................................................................................................... 151 

3  Finite element selection ......................................................................................................... 151 

4  Initial residual stress .............................................................................................................. 151 

5  Type of the analysis ............................................................................................................... 151 

6  Stepwise guiding procedure for the numerical simulation ..................................................... 151 

6

Summary

For the steel construction industry the research and development works performed refer mostly to conceptual design and calculation rules and their improvement. However there is a lack of development and optimisation of manufacturing techniques and its quality improvement of steel construction elements.

An important aspect to this is the control of geometry distortion and tolerance deviation due to welding and other manufacturing processes. In case of having a final geometry not fulfilling allowable tolerances one of the most common means to achieve realignment is flame straightening. Flame straightening is a local reheating with a torch. Due to the obstruction of the heat expansion an upsetting occurs. During cooling in the squashed zone shrinkage occurs that is associated with tension forces which lead to the desired deformation.

This process is practised mostly in workshops, but sometimes on site mostly with the intention to induce a defined deformation onto the steel element. The application range of this technique is large. In the structural steel market, flame straightening is applied to finely adjust elements in view to subsequently connect them by welding or bolting. It may be also applied to repair damaged elements, for instance the main girder of bridges submitted to an impact load by oversize vehicles. The technique may allow repairing the bridge without dismantling the whole structure. Similarly the flame straightening may be applied for railway applications. The railway companies use light profiles (IPE, channels and angles) as structural elements for rolling stock materials. These products need frequent repair along the life of the wagon. Due to impacts, the steel pieces are effectively frequently deformed. A thermal procedure is applied by flame straightening. The thermal conditions are not precisely defined. These works are made in internal workshops of the railway companies but also in other workshops of sub-contractors or customers throughout Europe. Application may also be find in the steel production process itself. The heat straightening technique could be applied in the rolling mill as a possibility to correct certain geometric defects of profiles, which are hardly modified with the press. The "out of square" defect is characterized by a lack of flange parallelism on beams. The "curved head" defect is a non-linearity affecting the extremity of the beam. These defects are generated by incidents on the rolling process or heterogeneous cooling cycles after hot rolling.

In summary, not only the correcting of a geometrical imperfection may be pursued by flame straightening but also various other aims: facilitating the assembly, cambering or building, repairing of damaged products, overlaying of inner stresses etc..

The process of flame straightening suffers from several shortcomings that have strong impacts on the quality, safety and -very obviously- on the economy of the manufactured steel construction. These are:

1. The flame straightening process to achieve the desired geometry of a steel construction element is almost a matter of empirical craftsmen-knowledge and thus neither really explained nor quantified, although the physical background is highly scientific.

2. The procedure of flame straightening is only reliable for a few “standard-cases”. In many other cases the amount, location, sequence and holding time of heating is not clear. For these cases often several trials for searching the real parameter-set have to be performed until the expected geometry can be achieved. Obviously this “try-and-error” process reduces significantly the economy of steel constructions.

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3. Due to the heating of steel at higher temperatures with different holding times especially in cases of modern high strength steels (HSS) an exceeding of temperature and time may have detrimental effects on strength and toughness.

4. On the other hand the reduction of the yield strength respectively the proportionality limit at higher temperatures leads to the necessity to reach temperatures above a certain level to induce a straightening effect by the heating process.

The objective of the research project is to thoroughly clarify the material mechanisms of flame straightening, the transfer of results to complex steel members and details. Thus a significant improvement of the procedures, the efficiency and the feasibility of flame straightening can be achieved. The improved flame straightening-knowledge and improved –procedures results into a more precise geometry, less detrimental effects on the material, less energy input, quicker and healthier production and therefore leads to a better economic position avoiding expensive repair work. Research of flame straightening and its adequate parameters of high strength steels are of particular concern within this research project as almost no research results are available.

The research work is divided into six work packages which are interdependent. As a basis for the subsequent working parts in a first step a literature survey is performed in work package 1. In work packages 2, 3 and 4 the experimental investigations are included while work package 5 covers the numerical examinations of the project. In parallel to the experimental and numerical studies the derived results are used to explain the mechanisms during the flame straightening process by analytical means. In work package 6 a guideline is elaborated for the presentation and clarification of the flame straightening process.

Literature survey

As a basis for the subsequent working parts in a first step a literature survey is performed. More than 100 literature references were collected, analysed and summarized. The results of this survey are not presented as a chapter of this report but given as a separate document which can be downloaded from the project web page (link in Appendix 1). The relevant parts were integrated in the guideline and a list with a selection of references is given at the end of the report.

Experimental investigations

By a large set of small scale tests and large scale tests the thermal and mechanical behaviour of different structural components during the heating process is investigated. For the experimental investigations the most common steel grades S235, S355 and S460 in steel construction have been selected for the investigation by small scale and large scale tests. The high strength steel grades S690 and S890 have been investigated by small scale tests.

Small scale tests with plate specimen of two different thicknesses (20 mm, 50 mm) were performed in a line heat process. In a parametric study the process parameters, plate thicknesses, steel grades and clamping conditions have been varied. From the tests the effect of the heat input, related to the velocity of the torch, could be identified. Maximum temperatures, cooling times between 500° C and 300° C (t5/3) and the resulting out-of-plane deformations increase with increasing heat input per unit length. For the deformations depending on the plate thickness an upper limit value of the heat input exists, from which the deformations start to decrease due to a reduced thermal gradient over the thickness and starting phase transformations. The value of this limit depends mainly on the plate thickness and is located at higher values for larger thicknesses. A second effect of the plate thickness is that for larger thicknesses the heat conduction tends to be more three dimensionally which leads to smaller cooling cycles t5/3 or an increase in the required heat input.

From the small scale tests the effect of the material strength is clearly identified. In the free end test setup the resulting deformations decrease with increasing strength; in the fixed end tests with increasing strength the remaining internal forces in the specimen increased.

Residual stress measurements with ultrasonic and sectioning method on the small scale specimens showed that the residual stress distribution arising from the flame straightening process resembles those of a welding process. Due to the more distributed heat flux distribution of the gas nozzles the region subjected to tensile stresses is wider. The maximum tensile stresses reach nearly the yield strength of the base material for each material investigated. For specimen with higher heat input and therefore

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higher maximum temperatures the tensile stresses in the centre tend to compression due to the starting in the phase transformations during the heating phase.

The material investigations after the flame straightening processes with the highest temperatures show no significant changes in the mechanical behaviour with one exception (S235J0, t = 20 mm), where the material toughness is reduced. For the high strength steels (S690QL, S890QL) the bainitic tempered microstructure is partially lost in the heated area, transforming into a ferritic and perlitic formation with fine grain for both investigated thicknesses. Nevertheless, the mechanical testing of these steels shows no significant changes.

Additional tests have been performed with small plates heated in a furnace to simulate a penetrative heating process like e.g. in a vee heat of a beam. The temperature cycle of a penetrative heating has been successfully simulated. The subsequent mechanical testing led to the conclusion that limiting temperatures stated in CEN/TR 10347 [3] should be observed for penetrative heating patterns.

Beside these parametric investigations by additional tests procedures have been developed to quantify the heat input of the nozzles used for the small and large scale tests and to identify the heat flux distribution for the numerical analysis.

To transfer the knowledge gained by the small scale tests a large number of large scale trials on complex structures has been performed in the steel construction workshops. The adjustment and handling of the torch is done by skilled operators of the workshop. Another objective of the trials is to prove that flame straightening practiced at recommended temperature chosen in accordance with the CEN/TR 10347 [3] was effective to generate a deformation or to straighten a distortion. The tests have been divided into four parts according to their configuration.

In the first part tests on profiles to generate a bending along the strong and the weak axis on both rolled and welded sections (I, H, T and L shapes) have been performed. For the bending of beams either continuous (e.g. heat band on the flange) or discrete heating patterns (vee heats or banded heats) proved to be applicable. Continuous heating patterns lead to a continuous curvature while discrete heating patterns lead to a polygonal curvature. From the results of the tests it can be concluded that the location of the heat pattern has a significant influence on the resulting deformation. The localized heating patterns (like vee heat) proved to be effective when located near additional loads due to the additional compressive stresses at the location of the heating pattern. For members with high moment of inertia (IPE, HEA) the heating results tend to zero without additional restraint and higher deformations are only achievable by additional loads or restraints. For profiles with small inertia relative large displacements were obtained for e.g. the T member about the y-axis without an additional load.

The effect of the material strength that was observed in the small scale test was also visible in the large scale tests. A comparison of the rolled and welded profiles showed no clear tendencies. For bending about the strong axis the results for welded and rolled sections were approximately the same while for bending about the weak axis the results of welded were about 70% larger than those of the rolled sections. In general instabilities might be possible when using additional weights in a load controlled way which may result in an additional displacement or, in the worst case, lateral torsional buckling. For the practice the handling of additional weights should be secured (e.g. by a crane) or path controlled restraints should be chosen.

In part two plate elements were tested in three different configurations (without stiffeners, plates with longitudinal stiffeners and plates with longitudinal and transversal stiffeners) and two different steel grades. The intention of the tests was to straighten the deformations from the fabrication processes (welding) or to induce a defined bending or buckle. Like for the small plates and the beams from these tests the influence of the material strength was also visible. This applies for the deformations from both the welding and the flame straightening process. The line heat process proved to be successful. Instead of this the applied heat spot patterns showed small effect on the out-of-plane deformation of stiffened plates. Without an additional restraint a deformation could not be induced. By the use of a mechanical jack a deformation was generated where the heat spots near the mechanical jack had the largest effect.

Part three of the large scale testing programme covers the straightening of distortions form fabrication processes like welding and rolling. Several cases from the rolling mill have been examined with out-of-square-defects and curvatures about the strong and weak axis. For out-of-square defects of profiles line heats on the web led to the desired straightening effect. For the curvature about the strong axis instead

9

of discontinuous vee heat patterns a longitudinal line heat was applied and proved to be very effective in straightening the curvature of the beam.

Butt- and T-welds have been fabricated with varying steel grades, plate thicknesses and weld throat thicknesses. The resulting angular distortion was straightened by the application of line heats on the backside of the welds. Two parallel line heats were applied on the T-elements in several cycles to straighten the deformations from welding.

In order to induce compressive stresses at the points of possible crack formation at the end of beams different heating patterns have been applied in part four of the large scale tests. For free beam end, half end cover plate and cope cut detail heating patterns were applied in a defined distance from the probable crack location and subsequently, the resulting residual stresses have been measured using the ultrasonic method. It was shown that by heating processes compressive stresses can successfully be induced at hot spot locations for crack initiation.

The investigation of material properties on large scale specimens shows that for superficial heats at temperature ranges below 650-700ºC, no significant changes are expected in the mechanical behaviour of the materials. For penetrative heats with temperatures above 700ºC some minor changes in mechanical response were observed, except for S235 steel grades in which, as seen in small scale tests, some significant variations in toughness and ductility parameters can be expected. For all these reasons, it is recommended not to perform the flame straightening processes at temperatures above 700º C for penetrative heating patterns, particularly for the case of ordinary ferritic-perlitic steels. For short superficial heating, higher temperature could be applied according to CEN/TR 10347 [3].

A general observation of all tests is that due to the manual processing the heating in the workshop is a very probabilistic process. Parameters like the adjustment of the torch, the distance between torch and element, the velocity of the torch (e.g. in a line heat), the heating times (e.g. for a vee heat or heat spot) and the oscillation parameters depend on the skills and experience of the operator. Measurements of the heating times for the flange of a vee heat showed that the standard deviation is at least 10 % of the average value.

Numerical investigations

The structural elements have been simulated by FEM-techniques to analyse the time dependant thermal, metallurgical and mechanical processes in a way that is not possible only by experiments to get an inside view of the mechanisms during the flame straightening trials. The numeric models were calibrated by the small scale tests and the full scale trials. Three different FE-codes (ANSYS, ABAQUS and SYSWELD) were used to identify adequate simulation techniques for the flame straightening process.

By numerical simulations with SYSWELD plates have been simulated to extend the knowledge gained by the small scale tests. All tests with short holding time have been recalculated and the results showed good agreement with the experiments. The analysis of the line heat process showed that the formation of the plastic strains is a three-dimensional mechanism which can hardly be handled by analytical means. Therefore parametric studies have been performed to analyse the effect of the heat input, the plate thickness, the steel grade and the size and type of nozzle on the resulting temperature cycles and deformations.

The linear relationship between the heat input per unit length and the maximum surface temperature and the resulting deformation in terms of the rotation angle was confirmed also for other combinations of plate thickness and nozzle. Due to the lower surface temperatures with multi-flame rosebud nozzles a higher heat input per unit length can be applied than with single orifice nozzles and, therefore, the resulting rotation angles at the line heat can be increased. The limit value of the heat input per unit length leading to decreasing deformations with growing heat input, which was observed in the small scale tests, could be recalculated by numerical means also for other combinations of nozzle, plate thickness and steel grade.

By division of the heat input per unit length by the plate thickness and multiplication of the rotation angle the effect of the plate thickness can be eliminated and, therefore, the relation between heat input and deformation can be given by one line for one steel grade and type of heat flux distribution. By the results of the small scale tests this line could be normalized to one line for all steel grades. With this linear relationship it is possible to estimate the deformation of a line heat process on a plate structure.

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Coupled thermo-mechanical transient FE analyses were performed on T-elements for two different heating patterns, for 10 and 20 mm thick plates and for two different material grades. With the simulations the deformations of the T-elements investigated in the frame of the experimental investigations could be recalculated and the results showed good agreement with the deformations of the test specimen.

Numerical simulations on beam structures with ABAQUS were performed in order to evaluate the influence of initial residual stress, the magnitude of external load, the size of heating pattern and the heating sequence on the final displacement. In general, the introduction of the residual stress in the numerical model results in lower final displacement. The consideration of residual stresses reduces the influence of the heating sequence (web-flange vs. flange-web). For vee heats with a large width and by application of high external loads the formation of a local plastic hinge at the location of the vee heat is possible and should be considered, because out-of-plane deformations of the flange and the web are possible especially for slender profiles. Up to the formation of the plastic hinge the deformation increases linearly with the applied external forces.

To derive prediction means for beam structures parametric studies have been performed with a component method. The piece of the beam surrounding the heating pattern is “virtually” cut out of the whole beam and the single vee heat is simulated using the flame model and process parameters of a realistic flame bending process. From the resulting deformation of this component the rotation of at the heating pattern is calculated. The deformation of the whole beam is calculated by superposition of the single deformations. These simulations proved that the relationship between the bending moment arising from external loads or restraints and the remaining rotation is linear, and that with increasing bending moment the deformation grows. This model has been calibrated by testing many real beams and the results obtained were in good agreement with them. Moreover, the fact that an entire beam can be simulated using just two or three single simulations leads to a significant reduction in the amount of time required to simulate it, constituting one of the main advantages of this method.

Analytical investigations

Based upon the experimental investigations and the numerical simulations the mechanisms of the heating patterns have been further investigated in the view of the prediction of the deformations of bar shaped structures. Two analytical models have been derived. With model 1 critical temperatures Tcrit,1 and Tcrit,2 can be calculated with which the beginning plastification in the colder part of the section can be set limits. The resulting deformations can be estimated by assumption that the complete thermal strains result in a plastic shortening of the vee heat and therefore in a rotation angle. By model 2 the plastic strains are derived using a fibre model of the section at different locations of the heating patterns. Assuming the hypothesis of Bernoulli and linear elastic – ideal plastic material behaviour the strain distribution is calculated in this way that the resulting stress distribution fulfils the equilibrium with the internal forces at the location of the heating pattern. Compared to the experimental results the analytical models give a good prediction when detailed information on geometry and sequence of the heating pattern and the temperature distributions are given.

The application of the derived analytical models is simple to use and quick possibility to predict the deformations from heating patterns on beam structures providing a relative well estimation for the critical temperatures and deformations. The application of the methods is presented in the guideline where also the procedures and means how to achieve the (critical) temperatures is shown.

Guideline

A guideline has been elaborated for the presentation and clarification of the flame straightening process. The guideline is addressed to people in the workshop but also to engineers working in the process engineering. The information gained through the literature review and the work within the work packages of the project has been summarized in a guideline giving an overview on the flame straightening process and on different kinds of heating processes.

The guideline covers all relevant fields for the application of the flame straightening process like explanation of mechanism, necessary equipment, heating patterns, sequences and restraints, procedure, adjustment of energy, quantification and control of heat input, avoiding an impact on the material, prediction means, and quantification of displacement. It includes an explanation of the application of the prediction means for plate and beam structures for which examples were worked out.

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The guideline is attached to the report in Appendix 2.

Furthermore a small guideline has been prepared concerning the numerical simulation of the flame straightening process. The guideline is attached to the report in Appendix 3.

General conclusion

A flame straightening process is supposed to be successful, when two conditions are fulfilled.

1. No detrimental effect has taken place within the material resp. the mechanical properties of the material have not changed due to the heating process. This is supposed to be the case when a maximum allowable temperature is not exceeded. For this purpose a recommendation for the maximum temperature has been elaborated for the different steel grades and heating types (superficial and full section heating) as a result of the experimental investigations. The recommended values are given in chapter 6.3 of the guideline (Appendix 2).

2. The aspired bending has been applied or the distortion from prior fabrication processes has been straightened. This aspect is addressed by the developed prediction means given in chapter 7 of the guideline for the deformations from line heats and heating patterns on beams.

The crucial point of a successful estimation of the heating process is to limit the various input parameters which often depend on the manual adjustment and handling of the torch. To achieve a good prediction of the results of a flame straightening process the control of temperature, heat input and heating time is of major concern as these are the input parameters for the estimation of the deformation. Procedures for the quantification of the heat input by a torch, the control of the temperature for superficial and penetrative heating patterns have been developed. The enlarged knowledge has been summarized and together with the prediction methods described in a guideline.

Even if the quantification of the heat input and the control of the temperature as input parameters lead to a small additional effort, by the derived prediction methods the flame straightening process becomes more effective and less critical, as no time consuming “try-and-error”-procedures have to be performed and the risk of a detrimental effect on the material has been reduced either by temperature control or prediction. Thus, a significant improvement of the procedures, the efficiency and the feasibility of flame straightening was achieved.

Nevertheless not all heating patterns and structural configurations could be investigated in the project. But the methods and procedures developed and applied within the project can be extended to other heating patterns and more complex structures.

12

Scientific and technical description of the results

Comparison of initially planned activities and work accomplished

In work package 1 a literature review was performed on the available references given on the flame straightening process. More than 100 literature references were collected and analysed in the view of the project objectives with the issues:

- Basics - Preparation - Adjustment of energy and oxygen partition of the torch (type of heating, type of burner

nozzle, gas mixture, gas supply, pressure ratio) - Description and definition / Intention - Explanation of mechanisms - Features/Results - Impact on Material - Impact on Energy - Certainty of success e.g. according to heating colours, deformation, etc.

A list of the references was arranged on the project web page where the references are directly linked to their scanned electronic versions. The references were individually revised, extracting the relevant information available into a summary. The information is related to the topics mentioned above insofar as many of the articles deal only with a few aspects of the attached above. All the collected information was included in its corresponding section, according to the list of topics of importance. The information of these summaries was compiled and presented into the mid-term report. An additional version is available for download on the project web page (links see Appendix 1).

In the framework of work package 2 a survey on the stress/strain measurement procedures was produced. The testing methods were separated in non-destructive, semi-destructive and full-destructive. A description of the physical aspects, advantages and disadvantages and recommendations for the respective field of application and their limitations were given. The ultrasonic method and, as a reference, the sectioning method were selected for the investigation of the small scale and large test specimen. To precise the results and to reduce the influencing effect of the texture for each material of the small scale specimen the acoustoelastic coefficients have been experimentally identified by measurement of the running times of the longitudinal and transversal ultrasonic waves in tensile tests. Nevertheless the calibration of the ultrasonic method with the available equipment proved to be very time consuming and elaborate. But finally the method has been applied on selected test specimen of work package 3 and 4 to derive the residual stress states after the flame straightening process.

By small scale tests the principal correlations between the heat input and the resulting deformations and material changes have been investigated in work package 3. Plate material has been procured in the steel grades S235, S355, S460, S690 and S890 in two thicknesses 20mm and 50mm and additional material from beams in steel grades (S355J2, Histar 355 and Histar 460). Microstructural and mechanical characterization of the material was performed before the tests for all materials and after the flame straightening trials for selected test specimen. This characterisation consisted of chemical characterisation, tensile tests, microstructural observations, hardness measurements, impact tests and fracture toughness estimations.

Three types of small scale tests (partially heated with free ends, partially heated with fixed ends and completely heated) have been performed. By variation of the process parameters, plate thicknesses and steel grades the heat input (gas flow, velocity of the torch) has been linked to the gained results in terms of temperatures, cooling times and deformations and material impact. Residual stress investigations have been performed on a selection of the small scale test specimen by the sectioning method and the ultrasonic method to obtain the residual stress distributions of the applied process.

Beside the main tests trials (about 100 tests) have been performed to identify the correct process parameters (like holding time, distance of the torch, velocity of the torch, oscillation, etc.) and the size and distribution of the heat input for the investigated nozzles. The procedures which have been developed to derive these parameters are stated within the guideline.

A large set of tests have been performed in work package 4 with large scale specimen on beam and plate structures to enlarge and transfer the knowledge gained in work package 3. The tests have been

13

performed in the workshops of subcontracted steel construction companies. The performance of the process (adjustment and handling of the torch) has been done by skilled workers of the workshop. The test specimens have been equipped with measurement devices to record process parameters, thermal cycles and deformations.

The tests are divided into four types:

1. Long profiles (41 tests)

2. Plates (6 tests)

3. Distortions of attachment, profiles and singles plates (16 tests)

4. Ends of profiles (12 tests)

For each performed test data sheets have been prepared giving the detailed test setup, specimen geometry, material parameters, process parameters and test results in terms of temperatures, deformations and change in material. Like for the small scale tests for selected specimen residual stress measurements with the ultrasonic method have been executed to quantify the induced residual stress states.

In parallel to the experimental works in work package 5 the elements were simulated by FEM-techniques to analyse the time dependant thermal, metallurgical and mechanical processes in a way that is not possible only by experiments to get an inside view of the mechanisms during the flame straightening trials. The numeric models were calibrated by the small scale tests and the full scale trials. Three different FE-codes (ANSYS, ABAQUS and SYSWELD) were used to identify adequate simulation techniques for the flame straightening process. Based upon the numerical and analytical investigations prediction methods have been developed for the line heat process on plates and for continuous and discontinuous heating patterns on beams. The methods have been calibrated and validated by the experimental results of the small scale and large scale tests.

Based upon the experimental and numerical works of work package 3, 4 and 5 the flame straightening process has been investigated analytically and prediction methods have been derived for line heats on plate elements and for continuous or discontinuous heating patterns on beams. The derived analytical models allow a simple to use and quick possibility to predict the deformations from heating patterns on beam structures providing a relative well estimation for the critical temperatures and deformations. The application of the methods was integrated into the guideline where also the procedures and means how to achieve the (critical) temperatures are shown.

Based on the knowledge gained by the analytical, experimental and numerical investigations and the literature review in work package 6 a guideline has been elaborated which covers the explanation of mechanism, the procedure, the necessary equipment, the types of heating patterns, sequences and restraints, the adjustment of energy, the quantification and control of heat input, the avoidance of an impact on the material, the prediction of deformations, the quantification of displacement and the certainty of success. Procedures for the quantification of the heat input and the control of the temperature have been developed to quantify the input parameters for the prediction means. The guideline is addressed to people in the workshop but also to engineers working in the process engineering. It includes examples for the estimation of the flame straightening result by the derived prediction means for plate and beam structures.

In conclusion it can be said that all tasks addressed by the proposal have been finished.

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1 Introduction

1.1 Overview of the project

Shape control by local heating is often used in processing of structural steels to correct deformations which occur during production or welding of plates or beams. Experience has proven the value of this operation in many applications and for many different steel types.

The flame straightening of a structural steel component is based on local heating of a limited region to be shortened in combination with the hindrance of thermal expansion by the cold vicinity and/or additional restraints. The process itself and the resulting deformations are based on the generation of negative plastic strains at the location of the structural element. This may be demonstrated by the chart given in Figure 1-1.

Figure 1-1: Flow chart of the flame straightening process Figure 1-2: Interrelationship of thermodynamics, mechanics and metallurgy

The process parameters are selected according to the original geometry and shape of the structural element. The choice of type, location and sequence of the heating pattern, the heat input and the heating times is governed by the distortion to be straightened or deformation to be induced. From the selected process parameters results a transient temperature field in the region of the heating pattern. Of high importance for the resulting deformation is the generation of a thermal gradient over the thickness for superficial heating patterns like line heats or within the section between heated and non-heated part for penetrative heating patterns like vee heats. This gradient leads in combination with the material and structural configuration to an obstruction of the thermal strains arising from the temperature field. This obstruction results in the formation of negative plastic strains in the regions of high temperatures. By the application of adequate additional restraints or loads the formation of plastic strains can be enforced. Due to this plastic strain field in the cooling phase the designated deformation arises. The plastic strains and balancing elastic strains form a strain distribution resulting in residual stresses. This mechanism is schematically given in Figure 1-3 for a line heat process on a plate.

Beside the mechanical mechanism the temperature cycle of the heating can result in phase transformations depending of the maximum temperatures and holding times during heating and the temperature rates during cooling cycle. These changes in the material can have a major influence on the mechanical result in terms of stresses, strains and finally, deformations.

To clarify these mechanisms of the flame straightening processes the relationship of the process parameters and the resulting material changes, plastic strains, stresses and deformations has to be identified and quantified for the applied heating patterns. From the scientific point of view the fields of thermodynamics, metallurgy and mechanics are decoupled and only the relevant interrelations are investigated (see Figure 1-2).

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Figure 1-3: Mechanism of line heat process on a plate (schematic)

With this major objective several flame straightening processes have been investigated experimentally, numerically and analytically within the research project. This is reflected by the structure of this report which is divided into these three main chapters. The works are divided into six work packages which are interdependent. As a basis for the subsequent working parts in a first step a literature survey is performed in work package 1. The results of this survey are not presented as a chapter in the report but given as a separate document which can be downloaded from the project web page (link in Appendix 1). The relevant parts were integrated in the guideline and a list with a selection of references is given at the end of the report. In work packages 2, 3 and 4 the experimental investigations are included while work package 5 covers the numerical examinations of the project. In parallel to the experimental and numerical studies the derived results are used to explain and quantify the mechanisms during the flame straightening process by analytical means. In work package 6 a guideline is elaborated for the presentation and clarification of the flame straightening process. The work packages and the corresponding chapter in the report are given in Figure 1-4.

Figure 1-4: Work packages of the project and structure of the project

The guideline elaborated within work package 6 is given in Appendix 2.

In the following the objectives of the project are stated and subsequently, the results of the experimental, numerical and analytical investigations are presented and a short overview on the elaborated guideline is given. The main conclusions from each chapter are summarized in chapter 5

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1.2 Objectives of the project

The objective of the project OPTISTRAIGHT is to clarify the material mechanisms of flame straightening, the transfer of results to complex steel members and details. Thus a significant improvement of the procedures, the efficiency and the feasibility of flame straightening can be achieved.

This major objective can be divided into several objectives according to the work packages of the project.

Objective 1: A survey on the actual knowledge shall be given reflecting issues for existing flame straightening processes and for the metallurgical impact on the material.

Objective 2: A survey on the existing stress-/strain-measurement procedures shall be elaborated and adequate stress-/strain- measurement procedures are chosen.

Objective 3: Applying the measurement techniques of the previous work-package in WP 3 laboratory tests with small scale specimens will be examined varying the mechanical constraints, steel grade, temperature, holding time, thickness and heat penetration. The objectives of the tests are to gain correlations of the input parameters to the achieved results and by using metallurgical knowledge and structural mechanics to explain these correlations analytically and to transfer theoretically to complex structures.

Objective 4: The results gained within the small scale tests shall be transferred to complex structures to verify and further develop the correlations derived.

Objective 5: Numerical simulations shall be used to extend the analytical predictions means from the experimental part of the project. The numerical models for the different types of structures shall be generated and calibrated by means of the test results and complex structures and elements shall be simulated.

Objective 6: Based on the knowledge acquired in the analytical, experimental and numerical investigations guidelines and recommendations shall be elaborated which clarify and simplify the flame straightening process and avoid an impact on the material.

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2 Experimental investigations

In the frame of the project a large set of trials were carried out with different objectives. In a first step small scale tests with different configurations have been performed to identify the principal correlations between the process parameters and the gained results in terms of impact on the material and deformations.

In a second step with tests on large scale specimen with different configurations the knowledge is transferred to more complex structures like beams and welded plate structures.

In the following an overview is given on the performed tests and the gained results.

2.1 Tests with small scale specimen

Three types of small scale tests have been performed. With the first and second type tests have been carried out with plates and a line heat process. With a third type of test the heating process of a penetrative heating pattern is simulated with small steel samples which are heated in a furnace.

2.1.1 Tests on partly heated specimen

In this part the principal correlations between the heat input parameters and the resulting deformations have been derived by plate specimen (length: 570mm, width: 570mm, thickness: 20 mm) on whom a line heat was applied. By variation of the velocity of the torch the influence of heat input per unit length has been investigated for two different plate thicknesses (t = 20 mm and t = 50 mm). The resulting temperature and deformation fields and the effect on the material have been examined for five different steel grades (S235JR, S355J2, S460M, S690Q and S890Q) and two different boundary conditions (free ends and fixed ends). Furthermore for the plate thickness t = 20 mm both short and long holding times have been investigated.

Beside these parametric investigations additional tests have been performed to quantify the heat input of the nozzles used for the small and large scale tests and to identify the heat flux distribution for the numerical analysis which are presented hereafter.

2.1.1.1 Quantification of heat input

2.1.1.2 Nozzles

For the processes of the line heat according to the agreed testing matrix nozzles with size N° 9 and N° 10 are used, which are shown in Figure 2-1.

Figure 2-1: Nozzles FB-A 8 to FB-A 10 (left side), detail of nozzle FB-A 9 (right side)

2.1.1.3 Tests to quantify the effective heat input of the torch

To quantify the effective heat input by a torch small scale tests have been performed using the following procedure. A test piece in form of a steel plate with dimensions 20x200x250 mm was heated by the selected torches in a line heat with a defined velocity and cooled directly after the heating in a defined mass of water. By thermocouples the increase of the temperature of the water has been measured. Parallel to this the gross heat input is quantified by measuring the gas flow. By application of the equations (2.1) or (2.2)

01 TTmcTmcQ waterwaterwaterwater in [J] (2.1)

with cwater = 4187 J/kg K and

19

t

Qq in [J/s] (2.2)

the effective heat input can directly been quantified using the measured temperature increase T. If the starting temperatures of steel plate and water are different, the temperature T0 has to be calculated by:

steelsteelwaterwater

steelsteelsteelwaterwaterwater

cmcm

TcmTcmT

,0,00

The results for the selected nozzles are given in Table 2-1. By variation of the distance of the nozzle to the surface of the specimen the optimum distance could be identified to 20 mm (see Figure 2-2).

Table 2-1: Heat input for nozzles N°9 and N°10

FB-A 9 FB-A 10

Effective heat input q

15.548 J/s 22.070 J/s

Total heat input qu

40.260 J/s 59.762 J/s

Degree of efficiency

38,6 % 36,9 %

Figure 2-2: effective heat input vs. distance of nozzle from plate surface

The results of further tests are given in chapter 3.1.2 of the numerical investigations. The detailed procedure is stated in the guideline (Appendix 2).

2.1.1.4 Tests to quantify the heat flux distribution

Beside the effective heat input for the numeric simulations also the parameters of the heat source distribution are necessary. To quantify these parameters resp. verify the parameters given by Rykalin [1] small scale tests have been performed in combination with numerical simulations acc. to WP 5 to derive these values.

A test piece in form of a steel plate (dimensions 20x200x250 mm) is heated with a line pattern along its longitudinal axis. By thermocouples the temperature cycles at different locations perpendicular to the heating line have been measured. Four thermocouples were positioned 1 mm below the surface and in distances 0, 10, 20 and 30 mm from the centre line of the torch. In a further step, the test has been recalculated by finite elements.

For the selected nozzles the following parameters were identified by calibration of the numerical parameters with the results of the experiments:

FB-A 9: qmax = 1,83 J/mm²s, r0,05 = 90 mm (V/t = 3000 l/h)

FB-A 10: qmax = 2,01 J/mm²s, r0,05 = 102,5 mm (V/t = 4500 l/h)

The resulting distribution of the heat flux is given in Figure 2-3 in comparison to nozzles N° 9 and N° 10 typically used for gas welding according to the literature [1].

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Figure 2-3: Distribution of heat flux for selected nozzles

2.1.1.5 Test setup

The test setup is mounted on a mobile testing frame. The test specimen is positioned horizontally in a cut-out of a steel plate with which the measuring frame is joined. The torch is mounted into an oscillating device that is fixed on a carriage moving on linear running gear consisting of a linear guiding with work gear spindle driven by a geared motor. The test setup is shown on the left side of Figure 2-4.

Figure 2-4: Test setup

(left: view of thermo vision camera, right: position of thermo vision camera)

Four thermo-couples are placed in drilled holes of each test specimen. One thermo-couple is located at the starting point of the heating process 1 mm below the surface where the heating up takes place to control the maximum temperature at this point. Three additional thermo-couples are located in the centre line of the specimen in different positions below the surface to measure the temperatures and thermal gradients arising during the process (see Figure 2-6). Additionally the temperature fields are recorded by a thermo vision camera mounted on a mobile platform (Figure 2-4, right). To have a reference to the typical measurement devices used in the workshop pyrometers and temperature indicator sticks are used in random checks. To control the gross heat input the acetylene consumption is measured with a gas flow meter which allows also measuring the working pressure and the temperature of the gas. For the free end setup six deformation transducers are used: three to measure the out-of-plane deformations along the free edge of the specimen and three to measure the in-plane deformations. In the fixed end setup six cantilever beams are fixed to the specimen, three on each side. Opposite

Oscillating device 

Torch 

Test  specimen 

Linear running 

gear 

Thermo vision camera 

21

cantilever beams are linked to each other by two load cells, so that in total six load cells measure the forces arising during the tests. Figure 2-6 gives an overview over the location of main measurement devices of the free end test setup. For the fixed end test setup

A draw wire sensor is used to control the heading of the linear running gear. The working pressure for acetylene is set to 1,1 bar, for oxygen to 5,5 to 6,0 bar. The ratio between oxygen and acetylene O2/C2H2 is adjusted to 1,05 to 1,1 (slight oxygen excess). Figure 2-6 shows the test setup with fixed specimen.

Figure 2-5: left: Free end setup, right: location of thermocouples for both test setups

Figure 2-6: Fixed end setup

2.1.1.6 Calibration tests

Using additional material S355 with thicknesses t = 20 mm and 50 mm calibration tests have been performed to identify the required process parameters.

t = 20 mm, short holding time

The short holding time tests with plate thicknesses t = 20 mm are performed in a line pattern with a single pass using nozzle FB-A 9 and no oscillation. The necessary parameters for the process to reach the aimed temperature are the velocity of torch vtorch and the time for heating up theating.

t = 50 mm, short holding time

The short holding time tests with plate thicknesses t = 50 mm are also performed in a line pattern with a single pass, but using nozzle FB-A 10 with oscillation. The necessary parameters for the process to reach the aimed temperature are the velocity of torch vtorch, the time for heating up theating and the adjustment of the oscillator.

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t = 20 mm, long holding time

The long holding time tests with plate thicknesses t = 20 mm are performed in a line pattern with multiple passes using nozzle FB-A 9 and no oscillation. The necessary parameters for the process to reach the aimed temperature are the velocity of torch vtorch, the time for heating up theating and the number of passes.

To identify the required heating time at the starting point of the test a large number of tests have been performed varying the velocity of the torch and the heating time. The results of these tests can be seen from Figure 2-7. There is a nearly linear relation between the achieved maximum temperature and heating time. For different velocities of the torch these lines are parallel, only the axis intercept grows non-linearly with decreasing velocity.

Figure 2-7: Maximum temperature vs. heating time

By variation of the velocity of the torch the maximum temperature can be influenced. The maximum temperature shows a non-linear dependency from the velocity of the torch which fits well to a potential function (see Figure 2-8).

Figure 2-8: maximum temperature vs. velocity of the torch (left: t = 20 mm, right: t = 50 mm)

As it can be seen from Figure 2-9 the cooling time t5/3 between 500° C and 300° C decreases with increasing velocity of the torch for both plate thicknesses.

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Figure 2-9: cooling time t5/3 vs. velocity of the torch (left: t = 20 mm, right: t = 50 mm)

The resulting out-of-plane deformations show a clear dependency on the heat input per unit length qs = Q/vtorch, see Figure 2-10. For plate thickness t = 20 mm a maximum deformation is achieved at a heat input of 6000 J/mm. This effect is due to the fact that with increasing heat input the thermal expansion grows which is then restrained by the cold material in the vicinity of the heated area. By further increase of the heat input nearly the whole thickness of the plate is heated, so that the stiffness of the remaining colder part decreases so that restraint decreases also. For plate thickness t = 50 mm, where the heating is superficial, an increasing tendency between heat input and out-of-plane deformation could be identified.

Figure 2-10: out-of-plane deformation vs. heat input per unit length (left: t = 20 mm, right: t = 50 mm)

2.1.1.7 Tests t = 20 mm, free clamping conditions, short holding time

A typical temperature cycle of test with plates t = 20 mm and short holding time is given in Figure 2-11.

Figure 2-11: temperature cycle of test S355_t20_free_sh_T3 (left: heating, right: cooling)

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The achieved maximum temperatures and cooling times t5/3 showed a good agreement with those derived in the calibration tests. Typical for the 20 mm plates is the low temperature gradient along the thickness of the plate during the heating phase.

The resulting out-of-plane deformation can be seen in Figure 2-12.

Figure 2-12: out-of-plane deformation cycle of test S355_t20_free_sh_T3 (left: heating, right: cooling)

With positioning of the torch onto the test specimen the plate deforms in the positive direction because of the thermal expansion of the heated zone. With beginning movement of the torch the negative deformation starts and develops until the torch has reached the end of the test specimen. After a small positive deformation the out-of-plane deformation grows further during the cooling cycle of the plate until the final deformation is reached at ambient temperature.

Taking all tests with t = 20 mm and short holding time into account the resulting out-of-plane deformation at the end of the tests can be illustrated vs. the yield strength of the material at room temperature and the heat input per unit length. This illustration is shown in Figure 2-13.

Figure 2-13: out-of-plane deformation vs. yield strength and heat input

It can be seen that for lower steel grades (S235, S355) the deformations are larger and the results show a maximum value due to heating through thickness while for the higher steel grades (S460, S690, S890) the values are lower and the development of the deformations is increasing with growing heat input. Due to the fact that the higher steel grades were not tested at the same high temperatures than the lower steel grades the maximum point could not be identified.

2.1.1.8 Tests t = 50 mm, free clamping conditions, short holding time

A typical temperature cycle of test with plates t = 50 mm and short holding time is given in Figure 2-14.

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Figure 2-14: temperature cycle of test S235_t50_free_sh_T3 (left: heating, right: cooling)

The achieved maximum temperatures and cooling times t5/3 showed a good agreement with those derived in the calibration tests. For the 50 mm plates the temperature gradient along the thickness of the plate is much higher than for the 20 mm plates.

The resulting out-of-plane deformation cycles can be seen in Figure 2-15.

Figure 2-15: out-of-plane deformation cycle of test S235_t50_free_sh_T3 (left: heating, right: cooling)

The deformation cycle during heating and cooling resembles the cycle of the 20 mm plates. But the development of the deformations is smoother and the resulting deformations are smaller (up to 3 mm).

Taking all tests with t = 50 mm and short holding time into account the resulting out-of-plane deformation can be illustrated vs. the yield strength of the material at room temperature and the heat input per unit length. This illustration is shown in Figure 2-16.

Figure 2-16: out-of-plane deformation vs. yield strength and heat input

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In comparison to the 20 mm plates the development of the out-of-plane deformations with increasing heat input is increasing for the lower steel grades (S235, S355, S460). For the higher steel grades (S690, S890) the resulting deformations are small and show only a small dependency on the heat input.

2.1.1.9 Tests t = 20 mm, free clamping conditions, long holding time

A typical temperature cycle of test with plates t = 20 mm and long holding time is given in Figure 2-17.

Figure 2-17: temperature cycle of test S355_t20_free_lh_T3 (left: heating, right: cooling)

The aspired maximum temperature range is achieved by a multiple pass process with variation of the velocity of the torch. For the first cycles the velocity is slow in order to increase the temperature while for the higher cycles the velocity is larger in order to hold the temperature. The achieved maximum temperatures showed a good agreement with the aspired values.

The resulting out-of-plane deformation cycles can be seen in Figure 2-18.

Figure 2-18: out-of-plane deformation cycle of test S355_t20_free_lh_T3 (left: heating, right: cooling)

The deformation cycle during heating and cooling resembles the cycle of the short holding time. During the heating passes the deformation grows nearly linear with time. This tendency continues during the holding phase, but the heating passes and the interim cooling phases can be easily identified from the deformation graph. For higher temperatures the resulting deformations are very much larger than for the tests with short holding time.

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Figure 2-19: out-of-plane deformation vs. yield strength and maximum temperature

The main effects that can be seen from the test results are the effect of the maximum temperature and of the steel grade. For all steel grades the results show an increasing tendency with increasing temperature. For low temperatures of about 600° C the steel grade does not have a significant influence. But for higher temperatures (> 650° C) the dependency of the yield strength of the material becomes clearer.

2.1.1.10 Tests with fixed clamping conditions

The process parameters (type of nozzle, heat input per unit length, heating time) of these tests were selected identical to the free end tests to identify the influence of the clamping conditions. The measured temperature cycles were comparable to those of the free end specimen.

The measured forces of the load cells have been transformed into the normal forces in transversal direction and bending moments about the longitudinal axis within the specimen by the equations

, ,

, ∙ 240 , ∙ 490

These sums of all three load cell pairs are given in Figure 2-20 over the time for S235_t20_fixed_sh_T1.

Figure 2-20: development of normal force and bending moment in specimen S235_t20_fixeD_sh_T1 (left: heating, right: cooling)

The development of the normal forces and bending moments is affine to the deformation cycles of the free end specimen. Within the heating phase the forces and moments start in positive direction and invert the development after the heating up at the starting point has ended. During the rest of the heating phase the negative bending moment and compressive forces increase. Both reach a minimum at the end of the heating phase. During the cooling phase, forces and moments start to develop positively, but until

28

ambient temperatures are reached, there is another change of sign of the force and moment rates, so that a local maximum forms in the graph. An evaluation of the final values of all tests with thickness t = 20 mm leads to the graph given in Figure 2-21.

Figure 2-21: normal force and bending moment vs. yield strength and heat input per unit length (left: normal force, right: bending moment)

The affinity that was visible in the development of the internal forces during the tests is also observed in the final values of the forces and bending moments. There is also a clear influence of the heat input per unit length leading to a minimum value at about 4500 J/mm while both higher and lower heat input lead to higher values. The second effect comes from the yield strength. With increasing strength the resulting values of the bending moment and the normal force increase absolutely, so that for S690 and S890 the largest compressive forces and negative bending moments are reached.

2.1.2 Free ends completely heated

The design “free ends – completely heated” was foreseen to reproduce the flame straightening thermal cycle in a furnace. The thermal reheating cycles in the furnace must be carried out in to reproduce as well as possible the heating with the torch, which is particularised by a fast rise in temperature. The accurate control of the temperature will allow correlating the effect of the temperature on the mechanical properties. According to some of the large scale tests, the heat straightening along the strong axis might induce a drop of yield strength properties. This case was chosen for the simulation.

According to these experimental curves, as one example is illustrated at Figure 2-22b, the heating time to reach the limit of 700°C was very short, around 150 seconds. The cooling curves, which is linked to the penetration of the flame, were based on the one recorded on the case of the large scale test. A cooling time between 500°C and 300°C greater than 400 seconds must be generated to be representative of a penetrative heat. The Figure 2-22 shows a comparison between the thermal cycle in the furnace and with the real trial. The procedure applied in the furnace represents well the cycle obtained for a penetrative heat straightening.

a) b)

Figure 2-22: comparison of the thermal cycle: a) furnace simulation and b) penetrative heat straightening trial.

S355J2 - Thin thicknessesRise in temperature up to 750°C

Furnace at 1150°C

0

100

200

300

400

500

600

700

800

900

0 100 200 300 400 500 600 700 800 900 1000 1100 1200Temps [sec]

Te

mp

era

ture

[°C

]

Temperature one extremity Temperature other extremity

275 s 450s

0

100

200

300

400

500

600

700

800

0 100 200 300 400 500 600 700 800 900 1000 1100 1200 1300 1400 1500 1600 1700 1800 1900 2000

Time (sec)

Te

mp

era

ture

( °

C)

Heating

420

200 1/6 aile

T4 T20

t 5/3 = 455

(Pyrometer)

29

The samples used in this experiment were taken from sections. Two ranges of flange thickness were investigated: thin flange thicknesses (0-16mm) in S355J2 and higher thicknesses (40mm) in S355K2+M and S460ML. For the thin thickness, four reheating temperatures were chosen (750, 850, 950 and 1050°C). For this range of thickness (8-16mm) and for the different reheated temperatures, the cooling cycle between 500 and 300°C was estimated around 400 and 500 seconds and so associated to a penetrative heat for each case.

These trials show that for a penetrative heat pattern, a maximum temperature of 700°C must be respected. With the reheating temperature of 750°C and higher, the tensile properties are still satisfied but a decrease of yield strength and tensile strength up to 10% could be observed. For the thicker thickness (40mm), the impact of the reheating temperature of 650°C, 700°C, 750°C and 800°C were investigated on the mechanical properties. The cooling times between 500 and 300°C were around 700 seconds. The thermal cycle could be compared to a real heat straightening operations on a heavy section.

The results of the tensile tests of the samples reheated in a furnace at 650°C, 700°C and 750°C show that the properties are nearly identical to those of the base metal. For all these reheating variants the tensile properties prescribed are well respected for both grades.

On the other hand, this is not the case for the reheating of the samples in a furnace at 800°C where we note a decrease of the yield and tensile strength of about 50 to 60MPa for the S355K2 (a decrease of 15%)and of about 100MPa for S460ML (corresponds to a decrease of 20%). Figure 2-23 illustrates these results for the yield strength and tensile strength for the S355K2+M quality. The loss of mechanical properties indicates that it is not recommended to perform a penetrative heat straightening in this range of temperature.

Figure 2-23: Tensile properties (Re, Rm) after heat simulation in the furnace

This procedure which allows a very precise control of the thermal cycle demonstrates a reliable behaviour of the material in penetrative conditions up to maximum temperature of 750°C.

In real flame straightening cycle, the possibility of discrepancies between the intended maximum temperature and the reality must be integrated. It is assumed from WP 4 that differences of 50°C up to 100°C are realistic.

2.1.3 Material investigation on small scale specimen

To clarify how the heat straightening process affects the properties of common construction steel grades, a sequence of tests including tensile tests, Charpy tests, hardness measurements and microstructure analysis was developed comparing the material properties before and after applying the heating process. The steels selected for the study were the following: S235J0, S355J2, S460ML, S690QL and S890QL. The samples analysed were obtained from two groups of plates, with reference thicknesses of 20 mm and 50 mm, respectively. The 20 mm plates were heated up to a target maximum temperature of 800-850 ºC (long holding times) for S235, S355 and S460 steels and to a temperature of 650-700° C (long holding times) for S690 and S890 grades. The plates of 50 mm thickness were heated up to a target maximum temperature of 850-900º C (short holding times) for steels S235, S355 and S460 and to a temperature of 650-700° C (short holding times) for S690 and S890 grades.

30

The chemical composition of the base materials is given in Table 2-2. The mechanical properties of the base material are given in the correspondent graphs of the comparison between base and heated material.

Table 2-2: Chemical composition of plates (% mass) C Si Mn P S N Cu Mo Ni Cr V Nb Ti B Al

S235JR, t = 20 mm 0,160 0,220 1,04 0,017 0,017 <0,08 <0,03 <0,02 <0,08 <0,03 <0,01 - <0,03 - 0,03

S235JR, t = 50 mm

0,090 0,180 0,83 <0,01 0,020 <0,08 <0,03 <0,02 <0,08 <0,03 <0,01 - <0,03 - 0,03

S355J2+N, t = 20 mm

0,106 0,493 1,57 0,013 0,0008 0,004 0,032 0,015 0,058 0,045 0,001 0,027 0,004 0,0001 0,036

S355J2+N, t = 50 mm

0,145 0,487 1,58 0,012 0,0007 - 0,017 0,005 0,020 0,024 0,000 - - - 0,034

S460ML, t = 20 mm

0,108 0,436 1,60 0,016 0,0006 0,0052 0,036 0,011 0,050 0,052 0,002 0,027 0,003 - 0,036

S460ML, t = 50 mm

0,050 0,357 1,57 0,011 0,0003 0,0035 0,296 0,159 0,454 0,029 0,000 0,020 0,003 - 0,034

S690QL, t = 20 mm

0,170 0,298 1,30 0,011 0,0008 0,0049 0,022 0,102 0,030 0,034 0,001 0,023 0,005 0,0018 0,073

S690QL1, t = 50 mm

0,168 0,282 1,28 0,010 0,0007 0,0028 0,016 0,253 0,117 0,316 0,002 0,028 0,002 0,0019 0,080

S890QL, t = 20 mm

0,161 0,273 0,87 0,009 0,0010 0,005 0,033 0,504 0,510 0,471 0,044 0,012 0,002 0,0001 0,040

S890Q, t = 50 mm

0,153 0,305 1,42 0,017 0,0007 0,005 0,016 0,360 0,120 0,520 0,030 0,002 0,003 0,0018 0,072

2.1.3.1 Tensile Tests

Figure 2-24 represent the yield stress and tensile strength values of the material before and after the heating process.

Figure 2-24: Yield stress (left) and tensile strength (right) in the specimens obtained from the 20 mm thick plate

Figure 2-25: Yield stress (left) and tensile strength (right) in the specimens obtained from the 50 mm thick plate

YIELD STRENGTH FOR 20MM THICKNESS PLATE

0

100

200

300

400

500

600

700

800

900

1000

S235J0 S355J2 S460ML S690QL S890QL

Material Grade

Yie

ld s

tren

gth

(M

Pa)

Base Material

Heated Material

TENSILE STRENGTH FOR 20MM THICKNESS PLATE

0

100

200

300

400

500

600

700

800

900

1000

S235J0 S355J2 S460ML S690QL S890QL

Material Grade

Ten

sile

str

eng

th (

MP

a)

Base Material

Heated Material

YIELD STRENGTH FOR 50MM THICKNESS PLATE

0

100

200

300

400

500

600

700

800

900

1000

S235J0 S355J2 S460ML S690QL S890QL

Material Grade

Yie

ld s

tren

gth

(M

Pa)

Base Material

Heated Material

TENSILE STRENGTH FOR 50MM THICKNESS PLATE

0

100

200

300

400

500

600

700

800

900

1000

S235J0 S355J2 S460ML S690QL S890QL

Material Grade

Yie

ld s

tre

ng

th (

MP

a)

Base Material

Heated Material

31

As it can be seen in Figure 2-24 and Figure 2-25 neither the yield stress nor the tensile strength change significantly. The average percentage of variation for yield strength is below 7% and for the tensile stress below 4%. The 20 mm plate shows a slightly downward tendency whereas the 50 mm shows an upward tendency; nevertheless, in both cases the changes can be considered of minor importance.

2.1.3.2 Charpy impact tests

The test specimens for Charpy impact tests were obtained from the heated and non-heated parts of the plates and tested in a wide range of temperatures (from -80ºC to 250ºC). Transition curves were obtained for all steel grades and for the two analysed thickness. Figure 2-26, Figure 2-27, Figure 2.28, Figure 2-29 and Figure 2-30 show the energy vs. temperature curves for each steel grade (in blue in as-received state and in red after flame straightening).

Figure 2-26: Charpy curves (absorbed energy vs. temperature) for S235J0 steel.

Left, 20 mm plate; right, 50 mm plate.

Figure 2-27: Charpy curves (absorbed energy vs. temperature) for S355J2 steel.

Left, 20 mm plate; right, 50 mm plate.

Figure 2.28: Charpy curves (absorbed energy vs. temperature) for S460ML.

Left, 20 mm plate; right, 50 mm plate.

S235 20mm thickness plate

0

50

100

150

200

250

300

-100 -50 0 50 100 150 200 250Temperature (ºC)

En

erg

y (J

ule

s)

Base material

Heated material

S235 50mm thickness plate

0

50

100

150

200

250

300

-100 -50 0 50 100 150 200 250Temperature (ºC)

En

erg

y (J

ule

s)

Heated material

Base material

S355 20mm thickness plate

0

50

100

150

200

250

300

-100 -50 0 50 100 150 200 250Temperature (ºC)

En

erg

y (J

ule

s)

Base material

Heated material

S355 50mm thickness plate

0

50

100

150

200

250

300

-100 -50 0 50 100 150 200 250Temperature (ºC)

En

erg

y (Ju

les)

Base material

Heated material

S460 20mm thickness plate

0

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300

-100 -50 0 50 100 150 200 250Temperature (ºC)

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erg

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ule

s)

Base material

Heated material

S460 50mm thickness plate

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100

150

200

250

300

-100 -50 0 50 100 150 200 250Temperature (ºC)

En

erg

y (J

ule

s)

Base Material

Heated material

32

Figure 2-29: Charpy curves (absorbed energy vs. temperature) for S690QL steel grade. Left, 20 mm plate; right,

50 mm plate.

Figure 2-30: Charpy curves (absorbed energy vs. temperature) for S890QL steel. Left, 20 mm plate; right, 50 mm

plate.

Regarding the charpy impact values, as can be appreciated in figures above, each material behaves differently; the main features are:

- Concerning the 20 mm thickness plate, manufactured in S235 steel, a noticeable worsening in its properties can be appreciated. This fact can be caused by a precipitation of non-soluble solid state carbon into cementite. This kind of cementite is known as “tertiary cementite” and its quantity is almost negligible but when it is found precipitated in grain boundary, as can be appreciated in the microstructure (see Figure 2-31), it can affect the material toughness. For the 50 mm thickness plate the properties show a slight improvement after the heating process.

- In reference to the S355 steel grade, there is a slight improvement in its properties after the heating process for both thicknesses (Figure 2-27).

- The material obtained from the 20 mm thickness plate manufactured with steel S460 shows a substantial improvement in its properties after the heating process, especially in the ductile-fragile transition area. However the material from the 50 mm thickness plate manufactured with this same steel does not show any significant differences between the base material and the heated material (Figure 2.28).

- The material taken from the plates manufactured with the S690 steel shows a slight improvement after the heating process in 20 mm thickness. However, in 50 mm thick plate, the upper shelf limit is higher while the transition area and the lower shelf limit are reached at a higher temperature (Figure 2-29).

- The S890QL 20 mm plate shows a slightly upward tendency whereas the 50 mm plate shows an upper shelf limit and a transition temperature which is the same for both the base and the heated material; nevertheless, in both cases the changes are of minor importance (Figure 2-30).

As a general comment about Charpy impact behaviour, it can be concluded that, except for the case of 20 mm 235J0 steel, all the differences observed between as-received and heated materials fall into the uncertainty margin limits of this kind of tests.

S690 20mm thickness plate

0

50

100

150

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300

-100 0 100 200Temperature (ºC)

En

erg

y (J

ule

s)

Heated material

Base material

S690 50mm thickness plate

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Heated material

S890 20mm thickness plate

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erg

y (J

ule

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Heated material

S890 50mm thickness plate

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-100 -50 0 50 100 150 200 250Temperatura (ºC)

En

erg

y (J

ule

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Base material

Heated material

33

Figure 2-31: Micro structure for S235J0 taken from 20 mm thickness plate.

Left, as-received state; right, after flame straightening.

2.1.3.3 Micro Hardness tests

To determine the hardness (HV1) in the two steel plates, several measures were made from top to bottom of the thickness of the plate every 2 mm. The results in Figure 2-32 show the average value of the hardness through the thickness of the plate.

As can be appreciated in Figure 2-32, in all cases a noticeable decrease in the material´s hardness was obtained in all cases, this reduction being, in general, larger for the 20 mm thick plate specimens (22%, on average) than from those taken from the 50 mm thick plate (9.5%, on average).

Figure 2-32: Summary of the results of micro-hardness obtained for 20 mm (left) and 50 mm (right)

2.1.3.4 Microstructure observations

The five steels analysed experienced some changes in the microstructure as a consequence of the applied heating that justify their mechanical behaviour; in summary, the following significant features were appreciated:

- Concerning the ferritic-perlitic steels (S235J0, S355J2 and SA460ML) (Figure 2-33, Figure 2-34, Figure 2-35), a refinement of the grain can be clearly seen as well as a reduction in the thickness of the perlitic bands, which gives rise to a new microstructural arrangement for both thicknesses. Regarding the 20 mm thickness S235 a precipitation of “tertiary cementite” in grain boundary was detected (Figure 2-31).

- In high strength steels (S690QL, S890QL) (Figure 2-36, Figure 2-37) the bainitic tempered microstructure is partially lost in the heated area, transforming into a ferritic and perlitic formation with fine grain for both thicknesses.

Hardness comparison in 20mm thickness plate

0

50

100

150

200

250

300

350

400

S235 S355 S460 S690 S890

Vic

kers

HV

1

Base material

Heated Material

Hardness comparison in 50mm thickness plate

0

50

100

150

200

250

300

350

400

S235 S355 S460 S690 S890

Vic

ker

s H

V1

Base material

Heated Material

34

Figure 2-33: Microstructure for ferritic-perlitic steels (S235, S355 and S460) as-received material taken from 20

mm thickness.

Figure 2-34: Microstructure for ferritic-perlitic steels (S235, S355 and S460) heated material taken from 20 mm

thickness plate

Figure 2-35: Microstructure for ferritic-perlitic steels (S235, S355 and S460) as-received material taken from 50

mm thickness

Figure 2-36: Microstructure for ferritic-perlitic steels (S235, S355 and S460) heated material taken from 50 mm

thickness

35

Figure 2-37: Microstructure for high strength steels (S690 and S890) as-received material taken from 20 mm

thickness plate

Figure 2-38: Microstructure for high strength steels (S690 and S890) heated material taken from 20 mm thickness

plate

Figure 2-39: Microstructure for high strength steels (S690 and S890) as received material taken from 50 mm

thickness plate

Figure 2-40: Microstructure for high strength steels (S690 and S890) heated material taken from 20 mm thickness

plate

36

2.1.4 Residual stress measurement on small scale specimen

2.1.4.1 Measuring of residual stresses by sectioning method

The sectioning method is based on the principle that internal stresses are relieved by cutting the specimen strips of smaller cross section. The stress distribution over a cross section can be determined with reasonable accuracy by measuring the change in length of each strip and by applying Hook’s law. The analysis is further simplified by assuming that the transverse stresses are negligible, and the cutting process alone produces no appreciable strains. In practice, however, transverse stresses may exist, but the lower the transverse stresses are the more accurate results will be. Residual stresses formed due to sawing alone depend, among many other factors, on the spacing of the saw cuts, the plate thickness and the speed of sawing. This can be avoided by introducing wire electric discharge or water jet cutting.

The residual stresses were measured in two plates that were previously treated by flame in a similar way than at straightening process by RWTH. The plate material was S355J2 with thickness 20 and 50 mm, respectively.

The distribution of strain gauges and the cutting lines is shown in Figure 2-41. The strain gauges were mounted to the upper and bottom side of the plate. The sectioning was done by water jet cutting (Figure 2-42). During the cutting the temperature did not exceed 50°C. The maximum temperature was reached only for a short duration of time due to friction between abrasive material and steel plate.

Figure 2-41: Distribution of strain gauges and cutting lines

The strains were measured in the longitudinal direction in the cross-section perpendicular to heating band. Four measurements of the strains were performed in the transverse direction in the centre of heating band.

Figure 2-42: Water jet cutting of 50 mm thick plate (left), plate after cutting

The measured residual stresses for 20 mm thick plate are shown in Figure 2-43 and for 50 mm plate in Figure 2-44. The average residual stress in the transverse direction were 86 and 64 MPa on the heated side for 20 and 50 mm thick plate, respectively. The stresses on the bottom side were 46 and 9 MPa for 20 and 50 mm thick plate, respectively.

37

Figure 2-43: Measured residual stresses for 20 mm thick plate

Figure 2-44: Measured residual stresses for 50 mm thick plate

2.1.4.2 Measuring of residual stresses by ultrasonic method

The ultrasonic method was used to measure the residual stresses within the small scale and large test specimen. By the measurement of the running times of longitudinal and transversal ultrasonic waves at the same location the stress state of the component can be quantified. Using an approach of continuum mechanics taking into account the change of the ultrasonic velocity by the stresses the stresses x and y can be calculated by the following equations:

,

,∙ ∙ ∙

, , ,

, ,∙ ∙ ∙

, , ,

, ,∙ ∙ ∙

longitudinal ultrasonic velocity

, transversal ultrasonic velocity, polarised in direction i (x or y)

A, B, C, D, F, H, K combinations of the elastic constants of second and third order

The index “0” indicates the stress-free state of the material. Assuming that the stresses perpendicular to the plate are zero ( 0 , the equations simplify to

, , ,

, ,∙ ∙

, , ,

, ,∙ ∙

The resulting stresses are averaged values of the stresses of the volume traversed by the ultrasonic waves (a cylinder of 12 mm diameter and thickness of the plate). The measurement of stress gradients over thickness is therefore not possible.

Figure 2-45 and Figure 2-46 show the results for tests with plate thickness t = 20 mm, free end specimen with short holding time and the highest heat input.

-220

-120

-20

80

180

280

380

-200 -100 0 100 200

Distance [mm]

Str

ess

[MP

a]Heated side

Symmerty

Bottom side

Symmerty

-220

-120

-20

80

180

280

380

-200 -100 0 100 200

Distance [mm]

Str

ess

[MP

a]

Heated side

Bottom side

38

Figure 2-45: residual longitudinal stress for test S235_t20_free_sh_T3

Figure 2-46: residual longitudinal stress for test S355_t20_free_sh_T3

In the heated region tensile stresses occur which reach nearly the yield strength of the base material. In the outer parts of the plates compressive stresses arise whose absolute values are lower than those of the tensile stresses. Within the heated region, there is a local minimum of the tensile stresses which tends to compression. This may have its reason in the phase transformation within this part as the temperatures are between 850° C and 900 °C. The qualitative distributions of the residual stresses are similar for the all investigated materials. The height of the tensile stresses is linked to the yield strength of the base material. From the measured residual stresses no clear tendency could be found in view of the investigated parameters.

2.2 Large scale tests

2.2.1 Introduction

In order to transfer the knowledge acquired with the small scale test, large scale trials on complex structures were performed. All the tests were performed in the workshop and were handled manually by skilled operators.

One objective was to prove that flame straightening practiced at recommended temperature, chosen in accordance with the CEN/TR 10347 [3] was effective to generate a deformation or to straighten a distortion.

The data collected during these trials were used simultaneously in the numerical models. For all the tests describe here after, a data sheet was created with the initial geometry of the test specimen, the test set-up, the equipment and the procedure as well as the results (temperatures, deformation, restraint forces, residual stresses, material impact).

In a first stage, 41 tests in order to generate a bending along the strong and the weak axis on profiles (on both rolled and welded section, on I, H, T and L shapes) were realized. A pre-trials campaign was realized in order to gain some knowledge on heat straightening operations.

In a second stage, six tests on plates were performed. Different configurations were tested: plates without stiffeners, plates with longitudinal stiffeners and plates with longitudinal and transversal stiffeners. Bendings along the longitudinal axis and out-of-plane deformations were generated.

In a third step, 4 trials in order to straighten distortions on sections (for instance, sections at the exit of the rolling mill or in a workshop due to wrong manufacturing or transportation) and 12 trials to straighten distortion induced by welding operations were performed.

In a last step 12 tests were performed to induce compressive stresses at critical locations of details prone to a cracking due to Liquid Metal Embrittlement.

39

2.2.2 Testing of long profiles

2.2.2.1 Heat straightening operations without mechanical means

Heating patterns

Three types of deformations were performed: a deformation along the strong axis, a deformation along the weak axis and flange parallelism deformation. The expected deformation was around 20 mm for the strong axis and 10 mm for the weak axis and flange parallelism. Thin (flange thickness lower than 16 mm) and thicker products (flange thickness of 40mm) in S355 and S460 were used during this preliminary study.

Deformation along the strong axis Three vee patterns were applied on a beam of 6 meters long.

Deformation along the weak axis Two bands per flange were symmetrically applied on a beam of 3 meter long

Flange parallelism Two kinds of pattern were realized on section of 1 meter length (Figure 2-47):

- To act on the flange parallelism deformation, a band was applied at the web-flange connection.

- For the deformation of a flange, a band was applied in the flange.

Figure 2-47: heat straightening operations to act on the flange parallelism

Flame parameters

The heating time performed varied with the kind of heating and is linked with the temperature on the surface. The follow-up of the temperatures was realized by means of optical pyrometers and thermocouples. The acetylene flow gas was in general around 5.000 liters/hours. The ratio oxygen on acetylene was around 1,5.

Resulting bending and impact on the properties

The tests show that no modification of mechanical properties was observed for a flame straightening up to 700°C. Above 700°C the test showed that the time spent at the maximum temperature also influence the final mechanical properties. This parameter determines the gradient measured in the heated zone and the cooling speed after heating. It allows distinguishing between a superficial heat (high gradient, high cooling speed) and a penetrative heat (small gradient, small cooling speed). In our test the cooling time between 500 and 300 (t5/3) was selected to quantify the penetration of the heat.

1 m

40

By analyzing the t5/3 and the surface temperature, the following criteria were established: a heat is considered as superficial (not penetrating) if the t5/3 is lower than 300 s. In practical terms this situation is observed when a band heat is applied either to correct the weak axis curvature or to act on the flange parallelism. In this case no significant effect on the properties is observed up to 950°C.

A heat is considered as penetrating if the t5/3 is higher than 300 s. This situation is observed with a vee heat for strong axis correction. In this configuration the mechanical properties are affected above 700°C, slightly for t5/3 between 300 and 400 seconds, significantly for t5/3 above 400 seconds.

These observations are fully in agreement with the CEN/TR 10347 (European Guidance for forming of structural steels) [3]. The adequacy of the recommendations for the choice of the maximum temperatures in function of the flame straightening process is thus confirmed by our trials. To be noticed that the European guide differentiate between "short full section heating" and "full section heating with longer holding time". These categories are similar to those based on the t5/3.

The Table 2-3 summarizes this information.

Table 2-3: Synthesis of the results of mechanical properties afterheat straightening

Heat type

Superficial

Penetrating –

short full section

Penetrating –

with longer holding time

T(5/3 <300s) 300<T(5/3)<400s T(5/3)>400s

Weak axis-

Flange parallelism

(band heat)

Strong axis for thickness below 20mm

(Vee heat)

Strong axis of thickness upper than 20mm

(Vee heat)

T°_max<700°C No effect No effect No effect

700<T°_max<950°C No effect 5% >10%

The trials showed that to generate a deflection (either with a penetrating or a superficial heat), the thermal actions was not sufficient. Increasing the flow rate was not a solution. The addition of mechanical means should be required. For the correction of flange parallelism it is not necessary to use a mechanical load. The cold zone around the heated zone is sufficient to make the deformation.

2.2.2.2 Heat straightening operations with mechanical means

The observations and the learning lessons of this preliminary study were applied on new trials.

42 trials were foreseen on both rolled and welded profiles with I, T and L shapes. The two first cases were not performed (HEM 400 – S235J0). As compensation, a preliminary study was added to the testing programme (reported in 2.2.2.1) in order to investigate the heating operations without mechanical means . A supplementary test was also performed (see data sheet 4.1-43) in order to apply the pattern foreseen by the numerical simulations and observed the correlation with the reality.

The description of the different trials and the results are described in the data sheets which can be downloaded from the project web page. Different patterns were applied in order to induce the deformation. The vee pattern was applied in most of the trials along the strong and the weak axis. The goal was to generate a bending of 20mm. The location of the vee heat with respect to the bending moment and the use of additional restraint were studied and are clearly reported in the data sheets. Amongst the tests, 2 trials with a band pattern in order to induce a deformation along the weak axis were also performed (data sheet 4.1.4 and 4.1.6). All the test specimens used have a length of 6 m and are supported as single span beam.

Flame parameters The different gas and working pressures used during this series of trials were summarized in the Table 2-4.

41

Table 2-4: Summary of the flame parameters and restraint load used for the trials 4.1-1 to 4.1-43

Restraint load Gas Pressure oxygene

Pressure Gas

ArcelorMittal 4.1-3 and 4.1-5 50 tons Tetrene unknown unknown 4.1-4 and 4.1-6 5 tons Tetrene unknown unknown 4.1-43 3 tons Propane 6 1.2

RWTH 4.1-13 to 4.1-18 4.1-27 to 4.1-42

7,5 tons Acetylene

5,5 / 6 1,1

UCAN 4.1-7 to 4.1-12 4.1-19 to 4.1.26

4 tons Propane

Temperature /displacement investigation: For each case, the temperature cycles were measured by the thermocouples which were positioned in the mid-thickness and at the subsurface (1 or 2 mm below the heat) in both the web and the flange. The temperature was recorded by a thermo vision camera for the trials performed by RWTH. To control the gross heat input, RWTH used a gas flow meter which allow measuring the working pressure and the temperature of the gas. Additionally, for the test performed by RWTH, the vertical deformation, the rotation about the longitudinal axis and horizontal deformation were measured with deformation transducers.

Heating pattern :

HE and IPE sections along the strong axis: vee pattern

The sequence of the vee pattern was constituted of:

- First, the top side of the flange (upper flange) was heated - then one side of the web from the top to the bottom - and finally the other side of the web, also from the top to the bottom.

Later, for several cases, the sequence was changed and the web was heated before the flange. The Figure 2-48 shows one example for this vee pattern.

Figure 2-48 : Example of location and sequence of heating pattern along the strong axis

Typically, the heating was divided into several steps of three of four heating patterns. After each heating step the specimen was cooled to nearly room temperature. A sequence without a restraint load was applied and confirmed that no deformation was generated for the H and I section.

HE and IPE along the weak axis: vee pattern and band pattern

The flange of the profile is heated from both sides with the shape of a vee (for IPE and HE profiles both flange) (Figure 2-49).

Figure 2-49: Example of location and sequence of heating pattern along the strong axis

A band pattern was also investigated for rolled section, as illustrated on Figure 2-50 . The flange tip was first heated from the extremity to the middle. In a second step, the heating continued from the other extremity to the middle of the beam. The two flanges tips were treated simultaneity by two operators. A restraint load of 5 tons was applied for the two trials.

42

Figure 2-50: Band pattern investigated to induce a bending along weak axis

L and T members along the strong and weak axis

For the L members(4 trials), the deformation along the strong and the weak axis were generated by realizing vee pattern but with variation of the height of the angle side. Using smaller vee patterns deformation along the weak axis was achieved. An example of heat pattern is given in Figure 2-51.

Figure 2-51: Location and sequence to induce a bending along the strong axis on L profiles

T members along the strong axis :

Only a transversal heat strip was applied on the flange of the T profiles (2 trials) (Figure 2-52).

Figure 2-52: Location and sequence to induce a bending along the strong axis on T profiles

T members along the weak axis Vee heats located on one side of the flange of the profile were performed (2 trials) (Figure 2-53).

Figure 2-53 Location and sequence to induce a bending along the weak axis on T profiles

Recorded deformation: The vertical deformation recorded during the tests showed a behaviour that is similar to the small scale tests of WP 3. During the heating phase the heated zone expands which results in a negative vertical deformation (upwards). In the subsequent cooling phase the vertical deformation develops in the opposite direction asymptotically towards the final deformation. In the Figure 2-54, the deformation development is given for the test 4.1-13. The test was performed in three steps indicated by the three blocks of acetylene consumption (FM). In the first step three vee heats were applied without an additional load on the beam. During the heating the beam deforms up to 10 mm in the mid-span location. After cooling a deformation of 1.3 mm was observed. During the second step with four vee heats and an additional load of 7,5 to the negative deformation decreases, but the resulting deformation after the cooling is about 10 mm. In the third step, the same behaviour was recorded.

43

Figure 2-54: record of vertical displacements for test 4.1-13

Based on the analysis of the resulting bending of the test, the following main conclusions can be drawn so far:

- The location of the heat pattern has an influence on the resulting deformation. The nearer the heat pattern is located to the middle of the beam the larger is the deformation that results from this pattern.

- Restraining by an additional load leads to higher deformations. For members with high - moment of inertia (IPE, HEA) the heating results tend to zero and higher deformations are only

achievable by additional loads or restraints considering the beam length of 6,0 m. Therefore relatively large displacements were obtained for the T member about the y-axis without an additional load.

- For most of the tests there is a clear effect of the material strength. At same restraint load and with the same heating process, the lower the steel grade the larger is the deformation.

- The comparison of rolled and welded profiles has shown two tendencies: o For IPE profiles about the strong axis the results for welded and rolled sections were

approximately the same. o For IPE profiles about the weak axis the results of welded were about 70% larger than

those of the rolled sections.

2.2.3 Testing of plate material

The objective of the tests with plate structures was either to straighten distortions from welding or to apply a defined deformation.

2.2.3.1 Tests on unstiffened plates

In tests 4.2-1 and 4.2-4 plates without longitudinal and transversal stiffeners have been subject of investigation. As the distortions from welding were minimal the objective of these tests was to induce a deformation in terms of a bending about the longitudinal axis which is comparable to that obtained from the tests in WP 3.

2.2.3.2 Test setup and heating patterns

For the tests the flanges at the plate’s side were used as support for the plate in longitudinal direction. The flanges themselves were supported at their ends.

The test specimens were equipped with transformation transducers in longitudinal and transversal direction to measure the out-of-plane deformations. With four thermo-couples the temperature-time-curves were recorded for one heat stripe. Additionally the temperatures were measured by thermo-vision camera, pyrometer and thermocrome sticks.

Three heat stripes were applied in longitudinal direction, one in the centre line and two 250 mm to the left and the right of the centre line. An overview over the location of deformation transducers and the heating pattern is given in Figure 2-55.

2.2.3.3 Results

Figure 2-56 shows the resulting deformation of test 4.2-1.

44

Figure 2-55: location of deformation transducers and heating pattern of test 4.2-1 and 4.2-4

Figure 2-56: resulting out-of-plane deformation for test 4.2-1

In both tests the heating pattern resulted in large deformations. In test 4.2-1 (S235) the vertex of the bending has a value of 120 mm. For test 4.2-4 (S460) the maximum out-of-plane deformation is 92 mm.

2.2.4 Tests on plates with longitudinal and transversal stiffeners

Two tests were carried out with steel grades S235 and S460 on plates stiffened by flanges and transversal stiffeners. The specimen are consisting of three plate fields with a = b = 1500 mm and a thickness of the plate of 10 mm.

The main objective of the tests was to obtain a defined out-of-plane deformation in the plate field by application of different heating patterns consisting of heat spots at different locations.

2.2.4.1 Test setup and heating patterns

The measurement equipment is the same as for the long profiles:

Deformation transducers to measure the out-of-plane deformations

Thermo-couples to measure the temperature cycles

Thermo-vision camera for temperature fields.

The location of the deformation transducers is exemplarily shown in Figure 2-57 for field 1 of the specimen 4.2-2.

Figure 2-57: Location of Deformation Transducers (DT), Field 1 Figure 2-58: Field 1of test 4.2-2 with lifting jack

The number and location of the heat spots is varied from field to field to study the influence of these parameters on the resulting deformation. The number and location of the heat spots can be taken from Figure 2-59 for test 4.2-2 and Figure 2-60 for test 4.2-5.

45

Figure 2-59 Heating patterns for the three fields of test 4.2-2

Figure 2-60: Heating patterns for the three fields of test 4.2-5

After first tests showed that an application of heat spots leads to no significant deformation additional restraints in terms of lifting jacks were used for an initial deformation (see Figure 2-58).

2.2.4.2 Test results

By the use of the lifting jacks it was possible to impose an out-of plane deformation on the plate. In Figure 2-61 the deformation record is given for field 1 of test 4.2-2.

Figure 2-61: Deformation record of field 1of test 4.2-2

In the first step (up to 3.000s of the record) no jacking was applied, the plate deforms upwards in the direction of heating. The subsequent cooling brings no resulting out-of-plane deformation. After the jack was installed and a deformation of 5 mm was imposed in the middle of the plate the heating pattern was repeated. Due to the heating the plate deforms downward up to a mid-span deformation of 17 mm which reduces during the cooling phase to value of 3.6 mm.

These tests were repeated in the other fields using the heat patterns in Figure 2-59 and Figure 2-60. The finally measured out-of-plane deformations for both test specimens are given in Table 2-5 and

Table 2-6.

Table 2-5: Deformations for test 4.2-2 (S235)

Field Number of heats Additional restraint

Deformation

1 3 Yes 3,6

2 4 Yes 4,0

3 20 Yes 5,9

46

Table 2-6: Deformations for test 4.2-5 (S460)

Field Number of heats

Additional restraint

Deformation

1 49 Yes 5,9

2 24 Yes 5,3

3 17 Yes 1,5

In all tests relatively small displacements were observed. The maximum value for both tests was 5.9 mm which means about 1/250. For S235 20 heat spots were necessary while for S460 more than twice as much heat spots had to be applied. Increasing the number of heating spots did lead to an increase of the displacement, but not proportional to the number of heating spots.

2.2.4.3 Test setup and heating patterns

The measurement equipment is the same as for the other plates:

Deformation transducers to measure the out-of-plane deformations

Thermo-couples to measure the temperature cycles

Thermo-vision camera for temperature fields.

Exemplarily the heating pattern and the location of the deformation transducers are given in Figure 2-62 for test specimen 4.2-3.

Figure 2-62: heating pattern of test 4.2-3

Typically the heat stripes are positioned near the web of the stiffeners on the opposite site of the weld that means two heat stripes per stiffener.

2.2.4.4 Test results

As an example for the results gained for this kind of test specimen the out-of-plane deformations of test 4.2-3 is shown in Figure 2-63.

For test 4.2-6 with steel grade S460 the procedure with two heat stripes per stiffener lead to acceptable straightening results. For test 4.2-3 the deformations from two heat stripes per web brought too large deformations (Figure 2-63, b: FS1 and FS2). Therefore it was decided to perform only one single heat stripe directly above the web for the remaining stiffeners (FS 3, 5 and 6). After this (Figure 2-63, c) the excessive deformation from FS1 and FS2 had to be straightened by a single heat stripe from below (FS 7, Figure 2-63, d).

47

a) After welding b) After heat stripe 1 and 2

c) After heat stripe 1 to 6 d) After completed process (FS7)

Figure 2-63: Out of plane deformation of test 4.2-3 (S235)

2.2.5 Testing of distortions of attachment, profiles and singles plates

2.2.5.1 Distortions on sections

4 sections presenting interesting defects were selected. For each case, the complete description of all the data are reported in the data sheets. Two out-of-square defects and two sections with a curvature were straightened.

Flame parameters:

For the 4 tests, the same equipment was used. A blowtorch Rhona2001 (n°14003712) with the nozzle D3 (n°14003236) was used. The working pressure for the oxygen gas was around 6 bars and the one for the propane gas was around 1.2 bars. The maximum opening for the tap was chosen in order to reproduce the same conditions in the two trials. The temperature evolution was measured with thermocouples and pyrometers.

A) Out-of-square defects Heat straightening operations were led on two cases of out-of-square: a HEA260 beam in S235J0 (data sheet 4-3.11) and a HZ1080MB beam in S430GP /S460 (data sheet 4.3-12).

For these two trials, the goal was to straighten the out-of-square defects without overtaking the 700°C.

Case 1 : out-of-square on one flange HEA260 in S235J0 – data sheet 4.3-11 Heating pattern

For the HEA 260, a non-constant out-of-square was measured on the upper flange 3-4. The flange 1-2 is straight and no action will be performed on it. A band pattern of 50-60 mm was applied in the web, centered at a distance of 65 mm from the exterior flange, as illustrated in Figure 2-64.

48

Figure 2-64: Heat pattern to straighten an out-of-square on one flange (data sheet 4.3-11)

The thermocouples inserted indicated peak of temperature at maximum 800°C. Short cooling time between 500 and 300°C were found, around 150 seconds. A gradient of less than 100°C between the subsurface and the mid-thickness was found (thickness of the web 8mm).

Resulting bending and impact on the mechanical properties

The improvement brought by the flame straightening operation with no mechanical support is relatively small (around 10mm). The mechanical properties were tested at 1/6th width of the flange and in the web (heated zone and non-heated zone). The mechanical properties after heat treatment are in agreement with the requirements and the heat treatment performed in these conditions did not deteriorate the mechanical properties.

Case 2: out-of-square on the two flanges of a HZ1080MB in S430GP – data sheet 4.3-12

Heating pattern

The two flanges were deformed with the same out-of-square, around 10mm. The beam was placed on two supports in the “I” shape. 2 band patterns (heat A and heat B) were applied in order to straighten the bar, at the exit of the curvature radius. The axis of the heat was located at ± 55 mm of the flange surface (Figure 2-65). The heat A was first performed and the heat B was applied in a second step. Mechanical jacks of 5 tons were installed each meter, installed in order to avoid the contraction of the flange 2-4 during the heat. A working pressure by mechanical jacks was applied.

Figure 2-65 : Heating pattern to straighten a similar out-of-square on the two flanges on a section

The heating speed for each heat was around 4minutes/meter. The temperatures during the heat were well respected (maximum peak around 700°C). The t5/3 was evaluated for the curve at 2 mm depth. The three delta show similar cooling time, around 150 seconds. This could be considered as a superficial heat.

Resulting bending and impact on the mechanical properties

Based on the measurement and on the observations, we can conclude that the flange 1-2 was straightened on the length on the section. The flange 3-4 was straightened on the 2.5m first meters but

49

on the other 2.5 meters the flange 3-4 was too much straightened and the out-of-square was reflected on the opposite side.

No important degradation of the mechanical properties is observed after heat straightening. A small decrease of the tensile strength in the web less than 5% is observed (decrease of 25 MPa). No impact on the yield strength is observed.

Two samples with a curvature were delivered by the plant: one presenting a curvature along the strong axis on the all length (data sheet 4.3-13) and a second, presenting a curved head along the weak axis (data sheet 4.3.14)

Curvature defects

Case 1: Curvature along the strong axis – HEM 340 in 275JR – data sheet 4.3-13

The maximum curvature was measured at the middle of the section, with a rope tightened between the two extremities and was around 47mm.

Heating pattern

With regards to the resulting bending obtained after the heat straightening sequence, three heating patterns were applied in order to straighten the curvature. For the first heating pattern, a median band of 60mm width at the mid-width of the exterior flange, on all the length of the beam was heated. A load of 3.2 tons was applied at the mid-length of the beam. The objective is to straighten the section without overtaken 700°C (in a conservative procedure). The heating speed was around 4 minutes per meter. This band pattern located at the middle of the width on the exterior flange allows a reduction of the curvature of 15mm.

In a second step, seven rectangular bands were applied in order to straighten the resulting curvature. The first heated pattern was applied at the middle of the beam. After, alternatively, the others bands were applied each meter. The Figure 2-66 illustrated the sequence of the heat.

The sequence of the rectangular pattern was constituted by:

- a transversal band on the exterior flange (view : Top flange /upper face) – width of 60mm

- continuously at this band, a rectangular band was performed in the inner side of the flange

- finally a band was performed on the web, in the same continuity.

Figure 2-66: description of the transversal bands pattern (second heat performed for trial 4.3-13)

This procedure allows reducing the curvature of 6 mm, which is a small gain in comparison with the median band on the exterior flange. It seems that the band applied at the mid-width on the exterior flange show a high potential to reduce the curvature along the strong axis, in comparison with the 7 rectangular bands pattern. So it was decided, in a third step, to reproduce a band pattern on the exterior

50

flange, with a higher width of 100mm. A higher temperature of 900°C was applied, which is higher than in the two trials.

In order to reach the temperature of 900°C, the heating speed was slower that in the first case (median band where the recommended temperature was 700°C) and was around 12 minutes /meter.

Impact on the mechanical properties

The results of the tensile tests after heating show that the values are in agreement with the requirements.

Case 2 : F9 – curved head along the weak axis- HEB300-S275J0 – data sheet 4.3-14

This sample presents a curved head on the last 2m50 with a maximal bending of 13mm.

Heating pattern

Two heating patterns were applied: First, the flanges tips were heated. The flange n°3 was first heated. Once this was performed the flange n°1 was heated without stopping between the two heats. This heat was considered as a superficial heat and we allowed the operators heating to a maximum temperature of 900°C. The control of the temperature was measured with the use of a pyrometer (Trotec TP6) and the thermocouples.

The recorded curves show that we observed high temperatures, even higher than the recommended maximum. Short cooling times t5/3 of about 150 seconds have been measured, which confirmed that this heat pattern is a superficial heat. With this operation, we gained only 2 mm and reduced the bending to 11mm.

Secondly, 7 bands were heated: this trial was performed in two steps (Figure 2-67): The heating pattern for these bands is composed of a triangle on the exterior and interior flanges n°1 and n°3 and a band in the web, in the continuity. The triangle height for the exterior flanges is around 150-160mm and the triangle height for the interior flange is around 110mm. A load of 3.2 tons was used. It was decided to heat up to a maximum temperature of 700° C in the flanges, as if it was a penetrating heat and up to 900°C in the web as a superficial heat.

Figure 2-67: Description of the second heating pattern (band pattern) applied for the trial 4.3-14

With this operation, we managed to reduce the bending to 0 mm.

2.2.5.2 Testing of T elements

12 straightening tests (2 more than foreseen) were performed on two types of welded elements/specimens. The goal of the tests was to straighten the distortions due to welding. Seven T shaped and five plates were fabricated by manual MAG welding (see Figure 2-68). Different combinations for the thickness of the plates (10mm or 20mm), the thickness of the weld (4, 7 or 10mm) and the steel qualities (S235J0 and S460ML) were investigated. The detailed information on the fabrication, the geometry, the testing and the results is given in the Datasheets from 4.3-1 to 4.3-10a.

51

Figure 2-68: Distortions induced by the welding for T and plates assemblies

In order to straighten welding distortion, line patterns were applied in the longitudinal direction. A single line pattern or two line patterns were studied. To achieve the desired deformations (straightening), up to 3 cycles of heating were required. The welding distortions were from 4.3 mm to 10.1 mm for T elements and from 0.9 to 10.1 for plate elements. There was no noticeable relationship between steel grade and the distortions.

An oxy-acetylene torch size 8 was handled manually. The pressure of acetylene was 0.95 bar and the oxygen was 9.2 bar. The pressure was manually calibrated by the operator according to the shape and colour of the flame.

Displacements and temperature were measured during the straightening process. Displacements were measured in four points using inductive displacement transducers (IDT). The temperature was measured in eight points by thermocouples, positioned in different depths of the plate.

The temperature at 1 mm below the surface never exceeded 650°C. Higher temperatures could be achieved only by slower velocity of the heating torch, which would also cause higher temperature over depth that would decrease the straightening effect.

It should be mentioned that for the plate specimen, the heat straightening operations induced a significant bending of the longitudinal direction, especially for 10 mm thick plates (up to L/50). An effort was made to fix the longitudinal distortions by transverse line heats. The plate elements showed to be very sensitive to line heating patterns. Therefore line heats are not the most appropriate heating pattern for straightening presented distortions on thin plates. Hence, the welding distortions of plate elements were not successfully straightened.

The welding distortions could be successfully straightened with a sub-surface temperature of 500-600°C, a localized and high gas flow is recommended. The objective is to achieve high gradient of temperature over the depth of the plate. Otherwise an external force would be needed to achieve the same result.

It should be mentioned that for the plate specimen, the heat straightening operations induced a significant bending of the longitudinal direction, especially for 10 mm thick plates (up to L/50). An effort was made to fix the longitudinal distortions by transverse line heats. The plate elements showed to be very sensitive to line heating patterns. Therefore line heats are not the most appropriate heating pattern for straightening presented distortions on thin plates. Hence, the welding distortions of plate elements were not successfully straightened.

52

Table 2-7: The test protocol and results of flame straightening of T elements (four measuring points for each T element)

2.2.6 Tests on ends of profiles

This test series had an intention totally different to the other experiments. The cope cut and half end cover plate are critical in view of Liquid Metal Embrittlement (LME) during the hot dip galvanizing process. Due to residual stresses from fabrication and transient stresses arising during the dipping process a risk for formation of cracks during the galvanizing process is given. In order to induce compressive stresses at the points of possible crack formation different heating patterns have been applied.

For all details two steel grades (S235J0 and S460M) and two types of profiles (IPE 550, HEB550) have been investigated. The following heating patterns have been examined:

- At the free end a line heat in transversal direction has been applied in a distance of 120 mm from the edge to induce compressive stresses at the edge.

- For the half end cover plate cracks occur 5-10 mm below the cover plate in longitudinal direction. To impose compressive stresses at the location below the cover plate a line heat of 170 mm length has been applied in transversal direction in a distance of 120 mm from the plate.

- The crack formation at the cope cut typically takes place at the corner of the cope. The crack initiates under an angle of 45°. In a distance of 141 mm from the corner point a heating spot is applied.

The detailed heating patterns are shown in Figure 2-69, Figure 2-70 and Figure 2-71. For these heating processes a nozzle type 8A (20-30 mm) was used with a medium acetylene consumption of 44,3 l/min. The resulting parameters of the temperature are given in Table 2-8.

Width of the heating band [mm] A [mm] 20 20Position relative to centre of the web B [mm] -40 40Time of heating[s] C [s] 190 213Displacement per cycle [mm] E [mm] 1.3 1.2 3.4 4.1 0.2

E [mm] 1.1 1.7 3.8 4.7 0.2Width of the heating band [mm] A [mm]Position relative to centre of the web B [mm] 0Time of heating[s] C [s]Displacement per cycle [mm] E [mm] 3.3 2.8 3.2 2.7 2.8

E [mm] 3.6 3.3 3.5 2.5 2.4Width of the heating band [mm] A [mm] 20 20Position relative to centre of the web B [mm] -85 85Time of heating[s] C [s] 102 140Displacement per cycle [mm] E [mm] 3.2 1.3 1.7

E [mm] 3.8 1.7 0.9Width of the heating band [mm] A [mm] 35Position relative to centre of the web B [mm] -38Time of heating[s] C [s] 146Displacement per cycle [mm] E [mm] 3.9

E [mm] 4.9Width of the heating band [mm] A [mm] 30 30 30Position relative to centre of the web B [mm] -40 -40 40Time of heating[s] C [s] 135 121 102Displacement per cycle [mm] E [mm] 4.0 4.0 3.6

E [mm] 4.2 4.1 4.2Width of the heating band [mm] A [mm] 35 35 35 35 35Position relative to centre of the web B [mm] -43 -43 43 -43 43Time of heating[s] C [s] 130 245 185 192 200Absolute displacement [mm] D [mm] 2.3 3.9 4.8 4.9 5.3

3.1 4.8 4.8 6.6 6.2Displacement per cycle [mm] E [mm] 2.3 1.6 2.5 1.0 0.5

E [mm] 3.1 1.7 2.1 1.8 1.4Width of the heating band [mm] A [mm] 50 50 50 50 50 50 50Position relative to centre of the web B [mm] -50 50 0 -50 50 -50 50Time of heating[s] C [s] 149 145 95 113 113 130 157Displacement per cycle [mm] E [mm] 2.6 3.4 1.1 1.2 1.5 0.7

E [mm] 3.0 3.6 1.3 0.9 1.1 0.7

2.72.32.7

Cycle

600

788

3542.51232.3

4.55.7

5.43040113

1085.1

3.74.735

37.5

0450

0.80.4

2231923.4

1200

243

120

4.1150

1200

300

146

1 2 3Specimen

T5

T6

T7

T1

T2

T3

T4

53

Table 2-8: Parameters of the temperature for ends of profiles

Free end Half end cover plate cope cut

Tmax, ° C t5/3, s Tmax, ° C t5/3, s Tmax, ° C t5/3, s

IPE 550 622-666 177-189 781-811 142-168 975-1009 69-74

HEB 550 672-738 201-212 734-753 164-176 822-913 51-64

             

Figure 2-69: Heating pattern on free ends Figure 2-70: Heating pattern on detail half end cover plate

Figure 2-71: Heating pattern on detail cope cut

To verify the effect of the heating processes on the possible crack initiation locations the residual stresses in the region of the heating were measured by the ultrasonic method. It was measured before and after the flame straightening process. The location of the measurement paths and direction of the stresses is given in Figure 2-69, Figure 2-70 and Figure 2-71. The results of the ultrasonic investigation are given in Figure 2-72 to Figure 2-75 for four selected specimen.

From the graphs it can be seen that for all investigated details the applied heating processes proved to be successful. The crack initiation locations are subjected to compressive stresses in the range from -100 MPa to -480 MPa. Within the heated area tensile stresses occur which reach the yield strength of the material for the free end details and lower values for the cope cut and half end cover plate. For the cope cut detail location of the heat spot should be put nearer to the corner of the cope to increase the height of the compressive stresses. A distance of 100 mm should be more appropriate for this. The location of the heating pattern of the other two details does not need an alteration.

54

Figure 2-72: Transversal residual stresses, free end detail, S235, IPE 550

Figure 2-73: Transversal residual stresses, free end detail, S460, IPE 550

Figure 2-74: Transversal residual stresses, cope cut, IPE 550

Figure 2-75: Transversal residual stresses, half end cover plate, S460, IPE 550

2.2.7 Material investigation on large scale tests

In order to complete the information provided by the material impact investigations in the small scale tests programme, tensile and Charpy impact samples were obtained from the large scale specimens indicated in Table 2-10, where the most relevant details of each one of these tests are summarised. The tests were performed on specimens machined from non-heated areas (base material) and from heated ones (heated material), with the aim of evaluating the impact of the process in steel properties. The specimens where taken, for the cases of profiles, from the normalized location, at 1/6 flange width. The results for all these tests (tensile and Charpy) are summarised in Table 2-11.

The chemical composition of the profiles is given in Table 2-9. The mechanical properties of the base material are given in the correspondent table of the comparison between base and heated material.

For the case of plates (4.2-1, 4.2-2, 4.2-4 and 4.2-5) and profiles (4.1-13 and 4.1-17), it can be observed that the results obtained confirm the observations done with small scale tests. No significant variations on yield stress or tensile strength were detected in any case and slight differences were only observed in impact properties. Just for the case of S235 grade, a reduction in ductility parameters was observed, maybe as a consequence of some kind of microstructural transformation as detected in the small scale tests programme.

For the case of sections, in those with an imposed curvature along the strong axis (4.1-3 and 4.1-5), a deterioration of the tensile properties was observed. A local decrease of the yield strength up to 60MPa is observed for the two steel qualities (S355K2+M and S460M). The tensile strength was less affected for the S355K2+M than for the S460M. For the S460M, a drop up to 60MPa was observed for the tensile strength. In spite of this reduction the specifications are respected.

It must be highlighted that the thermal curves recorded by the mean of thermocouples indicated that the maximum of 700°C was overtaken during the heat treatment. This was also confirmed with the results of the simulation of the heat straightening operations in a furnace. Indeed, the tensile results of the

55

samples reheated at 650°C up to 750°C show a conservation of the mechanical properties. On the contrary, for a reheating temperature of 800°C a decrease of 50MPa for the S355K2+M and of 100MPa for the S460M was observed.

Table 2-9: Chemical composition of investigated large scale specimen (wt-%)

Data sheet quality C Mn Si Al Nb V Ti CEV

4.1-43 S355M 0.09 1.06 .16 0.003 0.029 0.002 0.017 0.33

4.3-14 S275J0 0.06 1.04 0.2 0.002 0.002 0.035 0.001 0.29

4.3-13 S275JR 0.07 1.1 0.17 0.002 0.051 0.012 0.005 0.31

4.3-12 S430GP 0.08 0.93 0.21 0.002 0.007 0.002 0.0008 0.28

4.3-11 S235J0 0.10 0.58 0.17 0.002 0.002 0.002 0.0005 0.27

4.1-3 / 4.1-4 S355K2 0.09 1.04 0.17 0.015 0.011 0.003 0.019 0.31

4.1-5 / 4.1-6 S460M 0.09 1.35 0.21 0.011 0.012 0.058 0.016 0.37

4.1-13 S235J0 0.09 0.57 0.18 - - 0.003 - 0.27

4.1-17 S460ML 0.14 1.33 0.18 0.003 0.020 0.057 0.001 0.41

4.2-1 / 4.2-2 S235J0 0.16 1.04 0.22 0.030 - <0.01 <0.03 0.34

4.2-4 / 4.2-5 S460ML 0.11 1.60 0.44 0.036 0.027 0.002 0.003 0.39

Table 2-10: Large scale specimens selected for material impact investigation

Structure Type Steel grade

Maximum temperature (ºC)

Specimen taken from Data sheet

Sections HEM 400- Strong axis S355K2 725°C(pyrometer) 1/6 flange width 4.1.3

Sections HEM 400- Weak axis S355K2 790°C(pyrometer) 1/6 flange width 4.1.4

Sections HEM 400- Strong axis S460ML 770°C(pyrometer) 1/6 flange width 4.1.5

Sections HEM 400- Weak axis S460ML 725°C(pyrometer) 1/6 flange width 4.1.6

Plates

S235J0

1122.1 central plates

4.2.1

S460ML 4.2.4

Plates

S235J0

772.0 central plates

4.2.2

S460ML 4.2.5

Plates T element S235J0 < 650 heated zone 4.3-2

Plates T element S235J0 < 650 heated zone 4.3-3

Plates T element S460ML < 650 heated zone 4.3-5

Plates welded plate S460ML < 650 heated zone 4.3-10

Profiles IPE 450 Strong axis S235J0 921.4 1/6 flange width 4.1.13

Profiles IPE 450 Strong axis S460ML 986.7 1/6 flange width 4.1.17

56

Structure Type Steel grade

Maximum temperature (ºC)

Specimen taken from Data sheet

Sections HEA260 out-of-square S235J0 790 (thermocouple) 1/6 flange width 4.3.11

Sections HZ1080MB out-of-square S430GP 700 (thermocouple) 1/6 flange width 4.3.12

Sections HEM340 - curvature S275JR 900 (pyrometer) 1/6 flange width 4.3.43

In sections, with an imposed curvature along the weak axis (4.1.4 and 4.1.6), the tensile properties after heat straightening do not show any difference. The cooling cycles indicates very short cooling time between 500 and 300°C. It can be considered as a superficial heat. The target bending was achieved with a temperature of 700°C.

The impact of the flame straightening on the T-elements (4.3-1, 4.3-2, 4.3-5) and welded plates (4.3-10) was investigated by standard tensile tests and Charpy V tests, on the base material and on the heated material. The results of standard tensile test show that heating of the plates did not have any significant influence on the yield stress, ultimate tensile stress, ultimate strain and strain at fracture. For the Charpy specimen, there were no significant difference between the base and heated material.

For straightening of out-of-square defects, (tests 4.3-11 and 4.3-12), the heat is applied following a band pattern localized at the exit of the radius of the web-flange connexion. The mechanical properties were evaluated in the standardized location (1/6 flange width), and those values are the ones included in Table 2. For a controlling purpose, the behaviour of the material was analysed in the heated area. No degradation was observed in this case.

Table 2-11: Results of Tensile and Charpy test on material from large scale specimens

Data sheet/ structure/ steel grade fy (MPa) fu (Mpa) A5 (%) Charpy (Joules)

4.1.3 Section

S355K2

Base mat. 412/427 530/532 28.5/26.5 >264 J at -40°C

Heated mat. 346/425 522/537 25.0/24.0 237 at -20ºC

4.1.4 Section

S355K2

Base mat. 412/410 530/525 28.5/28.5 >264 J at -40°C

Heated mat. 412/403 526/526 29.0/28.5 231 at -40°C

4.1.5 Section

S460ML

Base mat. 553/552 657/659 19.5/21.5 >200J at -40°C

Heated mat. 476/490 586/601 17.0/16.5 186 at-40°C

4.1.6 Section

S460ML

Base mat. 553/553 657/653 19.5/21.5 >200J at -40°C

Heated mat. 553/537 640/643 23.0/21.5 211 at -40°C

4.2.1 Plate

S235J0

Base mat. 270 444 40.5 194 at 0ºC

Heated mat. 305 454 41.5 199 at 0ºC

4.2.4 Plate

S460ML

Base mat. 539 709 28.0 208 at -20ºC

Heated mat. 495 698 30.0 192 at -20ºC

4.2.2 Plate

S235J0

Base mat. 270 444 41.0 194 at 0ºC

Heated mat. 305 453.9 33.0 180 at 0ºC

4.2.5 Plate

S460ML

Base mat. 539 709 28.0 208 at -20ºC

Heated mat. 522 673 29.0 191 at -20ºC

4.3-2 Plate S235J0

Base mat. 249/249 404/403 35.5/41 247 at 0°C

Heated mat. 267/281 401/407 34.8/35.3 259 at 0°C

4.3-3 Plate S235J0

Base mat. 398/410 499/500 32.5/33.1 195 at 0°C

Heated mat. 374/391 474/498 29.4/32.8 208 at 0°C

57

Data sheet/ structure/ steel grade fy (MPa) fu (Mpa) A5 (%) Charpy (Joules)

4.3-5 Plate S460ML

Base mat. 496/497 588/585 25.1/26.9 220 at -50°C

Heated mat. 554/527 609/608 24.7/27.8 202 at -50°C

4.3-10 Plate S460ML

Base mat. 517/520 578/579 26.1/26.5 297 at -50°C

Heated mat. 512/504 573/557 26.9/27.5 294 at -50°C

4.1.13 Profile

S235J0

Base mat. 303 453 39.0 139 at 0ºC

Heated mat. 309 457 31.0 171 at 0ºC

4.1.17 Profile

S460ML

Base mat. 458 627 27.0 117 at -20ºC

Heated mat. 476 611 27.0 136 at -20ºC

4.3.11 Section

S235J0

Base mat. 323 446 30.0 86 at 0ºC

Heated mat. 326 450 32.0 84 at 0ºC

4.3.12 Section

S430GP

Base mat. 524 619 27.0 Not tested

Heated mat 521 624 27.0 Not tested

4.3.13 Section

S275JR

Base mat. 336 447 33.0 Not tested

Heated mat. 343 455 35.0 Not tested

Finally, for the case of straightening of a curvature defect (4.3-13), no significant variation on mechanical properties was detected. As a general conclusion, the investigation of material properties on large scale specimens reveals that for superficial heats at temperature ranges below 650-700ºC, no significant changes are expected in the mechanical behaviour of the materials. For penetrative heats with temperatures above 700ºC some minor changes in mechanical response were observed, except for S235 steel grades in which, as seen in small scale tests, some significant variations in toughness and ductility parameters can be expected. For all these reasons, it is recommended not to perform the flame straightening processes at temperatures above 700º C for penetrative heating patterns, particularly for the case of ordinary ferritic-perlitic steels.

For short superficial heating, higher temperature could be applied according CEN/TR 10347 [3]. Although the trials performed during this study prove that deformation could be achieved without overtaken 700°C.

2.2.8 Residual stress measurement on beams

By the ultrasonic method presented in chapter 0 measurements were performed on beams of the large scale tests. For tests 4.1-15 and test 4.1-16 the results of the longitudinal residual stresses after the flame bending process are presented hereafter. Due to the measurement with waves perpendicular to the surface the measurement in the crossing region of web and flange is not possible.

In test 4.1-15 an IPE 450 (S355J2) was investigated with vee heats about the strong axis.

Figure 2-76: Location of residual stress measurement for test 4.1-15

58

Figure 2-77: Residual stresses in the flanges of test 4.1-15 (IPE 450, S355J2, vee heat n° 11 about the strong

axis)

Figure 2-78: Residual stresses in the web of test 4.1-15 (IPE 450, S355J2, vee heat n° 11 about the strong axis)

In the web compressive stresses are present in the regions near the flanges. In the middle part the web is subjected to tensile stresses with a local minimum at the axis of the profile. In the flanges the stress state is asymmetric. In both flanges the right side shows small tensile stresses while on the left side compressive stresses occur, in the upper flange with values up to 75% of the yield strength. This may have a reason in the additional loading which led in some tests to a lateral movement of the profile and especially of the upper flange. Due to the reduced yield strength during the heating cycle a part of the upper flange plastified.

In test 4.1-16 an IPE 450 (S355J2) was investigated with vee heats about the weak axis.

Figure 2-79: Location of residual stress measurement for test 4.1-16

Figure 2-80: Residual stresses in the flange of test 4.1-16 (IPE 450, S355J2, vee heat n° 4 about the weak axis)

The stress state shows tensile stresses in the heated region. In the colder part of the flange compressive stresses arise in longitudinal direction.

59

3 Numerical investigations

In parallel to the experimental works the elements were simulated by FEM-techniques to analyse the time dependant thermal, metallurgical and mechanical processes in a way that is not possible only by experiments to get an inside view of the mechanisms during the flame straightening trials. The numeric models were calibrated by the small scale tests and the full scale trials. Three different FE-codes (ANSYS, ABAQUS and SYSWELD) were used to identify adequate simulation techniques for the flame straightening process.

After presentation of the numerical model of the flame the simulations of plate and beam structures are presented including the main results of the calculations and the derived prediction means.

3.1 Numerical model of the flame

Rykalin [4] provides a model for the calculation of the heating of metals by flames. The distribution of the heat flow density of a flame whose axis is perpendicular to the surface of the material is rotation-symmetric to the axis of the flame. The intensity varies with the distance r from the axis. The heat flow density has its maximum value q̂ on the flame axis and decreases with increasing distance from the axis. It can be assumed that the heat flow density of the flame is Gaussian distributed and can be written in the form:

2

ˆ' rkeqrq (3.1)

The correlation between the maximum value q̂ of the heat flow density and the effective heat input q is

k

qqqk

q ˆˆ (3.2)

where k is the concentration factor of the Gaussian distribution. The concentration factor is directly linked to the radius of the heat source r0,05 by

205,0

05,0

33

rk

kr (3.3)

The radius of the heat source represents the range in which 90% of the heat flow is situated. By insertion of (6.4) in (6.3) and subsequent insertion of (6.3) in (6.2) the function for the heat flow density results to

2

05,0

2

2

3

205,0

3 r

r

rk er

qe

kqrq

(3.4)

For simulation purposes it is useful to define this function within the x-y-coordinate system:

2

05,0

22

2

3

205,0

3,

r

yx

rk er

qe

kqyxq

(3.5)

In Figure 3-1 the distribution of the heat flow density is given.

Figure 3-1: Distribution of the heat flow density according to equation (3.5)

61

3.1.1 Heat input and heat flow density distribution parameters in the literature

In [1] Rykalin gives thermal specifications for common burner nozzles with sizes 2 to 8 for use with acetylene. Beside the acetylene consumption the effective heat input and the degree of efficiency are described. By the degree of efficiency u the effective heat input q is linked to the complete heat input of the flame qu, given by the formula

uu qq / . (3.6)

For the description of heat sources which are necessary for simulation purposes the maximum heat flow density q̂ and the nominal radius of the heating spot r0,05 are given. Table 3-1 shows the thermal specifications, added by maximum heat flow density for the 3D-heat source model.

Table 3-1 : Thermal specifications for common burner nozzles

Burner nozzle

Consumption

C2H2

Effective heat input

q

Degree of efficiency

u

Maximum heat flow density

q̂

Nominal radius of heating spot

r0,05

l/h J/s % J/(mm²s) mm

2 150 1740 86,60 2,20 27,2

3 300 2618 65,13 2,72 30,1

4 500 3537 52,79 3,18 32,4

5 750 4490 44,69 3,60 34,4

6 1150 5776 37,49 4,10 36,6

7 1700 7270 31,92 4,63 38,8

8 2500 9124 27,24 5,21 41,1

9* 4000 12034 22,45 6,02 44,0

10* 7500 17425 17,34 7,29 48,2

* extrapolated

For steel plates with thicknesses greater than 20 mm sizes 9 or 10 are often used for flame straightening purposes. To derive the specifications for these burner nozzles the correlation between the parameter and the acetylene consumption is investigated for the known data and then extrapolated to the values for size 9 and 10. Figure 3-2 and Figure 3-3 show the correlations and Table 3-1 gives in the last two rows the extrapolated values.

Figure 3-2: Effective heat input and degree of efficiency vs. acetylene consumption

Figure 3-3: Maximum heat flow density and effective diameter of heat source vs. acetylene consumption

With growing acetylene consumption the effective heat input increases. Because of the small degree of efficiency for larger nozzles the growth is non-linear. Both maximum heat flow density and effective diameter of the heat source increase with growing acetylene consumption.

62

3.1.2 Experimentally derived heat input

Within the project several nozzles have been investigated with regard to the heat input parameters. The testing procedure given in chapter 2.1.1.3 was applied to identify these parameters. An overview of the parameters is given in Table 3-2.

Table 3-2 : heat input parameters for nozzles investigated within the project

Nozzle Type of nozzle Effective

heat input q

Total heat input

qu

Degree of efficiency

C2H2 gas flow

J/s J/s - l/h

FB-A 9 multi-flame rosebud 15.548 40.260 38,60% 3.018

FB-A 10 multi-flame rosebud 22.070 59.762 36,90% 4.480

N° 8 single flame 14.949 25.345 58,98% 1.900

Geyer A10 multi-flame rosebud 18.738 43.414 43,11% 3.255

7A single flame 11.550 28.941 39,94% 2.170

8A multi-flame rosebud 15.086 36.228 41,64% 2.716

From the results it can be seen that the degree of efficiency of the multi-flame rosebud nozzle is higher than those of single flame nozzles. Furthermore, the efficiency values for the both single flame nozzles are higher than those given in the literature (see Table 3-1). The reason for this discrepancy could not be identified.

3.2 Plate structures

3.2.1 The selection of an efficient finite element

The numerical modelling of the flame straightening process is an expensive task. Therefore, choosing an efficient finite element presents a unique challenge. The comparison of different solid and shell elements is presented in the sequel. The study was performed in Abaqus v6.7 software.

Figure 3-4: Temperature over depth of the plate for different finite element types at t = 39 s

A temperature gradient over the thickness of the steel element enforces a stress gradient over thickness. The study of finite elements was done on a numerical model with the same or similar mesh size. Steel plate b/h/t = 500/600/20 mm was heated with heating torch no. 8 according to Rykalin [1]. The flame of the heating torch was modelled by gap radiation option, where radiation was distributed according to Rykalin equation [1]. The flame was travelling with the speed of 8 mm/s in longitudinal direction of the steel plate (h = 600 mm). Figure 3-4 and Figure 3-5 show temperature at time 39 s and stress after cooling over depth of the plate in the intersection of symmetry planes of the plate. Stresses shown in Figure 3-5 are extrapolated from integration points to nodes and averaged for each node.

Temperature [°C] at t = 39 s

0

200

400

600

800

0 5 10 15 20Depth [mm]

Tem

p.

[°C

]

S4RT C3D8T

C3D8RT C3D4TC3D10MT C3D20T

C3D20RT

63

The first option was C3D8RT elements (8-node tri-linear displacement and temperature bricks with reduced integration). Since these elements have only one integration point, the stresses are constant over element. Full integration does not give significantly different results. The second option was C3D20RT (20-node tri-quadratic displacement, tri-linear temperature bricks with reduced integration). The problem with these elements is that temperature is active degree of freedom only in corner nodes. This is probably the reason why surface heat flux generated by radiation differs significantly from the analytically calculated heat flux. Due to this fact, temperature in Figure 3-4 is much lower than for other elements. Therefore, these elements should be avoided in these kinds of problems. Full integration gives very similar results. Elements C3D10MT (10-node modified displacement and temperature tetrahedron) have temperature as active degree of freedom also in mid-side nodes. C3D4T are linear version of previously mentioned elements (4-node linear displacement and temperature tetrahedron). The results of 10-node and 4-node tetrahedron give similar results as 8-node bricks (Figure 3-4, Figure 3-5). Shell elements are much more economic. Elements S4RT are 4-node doubly curved shell elements with reduced integration (finite membrane strains, bilinear temperature). Ten integration points over depth were used to calculate stresses in Figure 3-4, Figure 3-5. The results are very similar to those of 10-node tetrahedron.

Figure 3-5: Stresses over depth of the plate for different finite element types

It can be concluded that solid elements are too expensive for the studied problem. Shell elements S4RT are very efficient and will be applied in the numerical simulations with ABAQUS in the framework of this work package. The most effective solid finite elements are 10-node tetrahedron C3D10MT. These elements will be used when dealing with problems where modelling with solid elements is necessary.

3.2.2 Numerical simulations of plate elements

Numerical simulation of the small scale test specimens were performed to calibrate and validate the numerical models be the experimental results and to understand the process of formation of plastic strains more clearly to identify the mechanisms of the resulting deformations. The simulations were realized with the FE-Software SYSWELD which is, beside the thermal and mechanical part, also capable of covering the metallurgical changes which might occur und which may have an influence on the flame straightening result.

The numerical models were designed with solid elements with 8- respectively 20- nodes. To reproduce the interaction of the heated plates with the environment by convection and radiation the surface was meshed with 2D-elements. The flame is implemented using the numerical model of a two-dimensional heat flux presented in chapter 3.1.

The calculation process is divided into two parts. In a first step a transient thermal-metallurgical calculation is performed. The results of this calculation is the temperature field T(x,y,z,t) and the field of the phase proportions pi(x,y,z,t) for each time step of the calculation. Based on these results the subsequent mechanical calculation is performed using the phase proportions to derive the changed material parameters and the temperatures to derive the thermal strains th. The process is typically divided into two phases. First, the flame is positioned at the starting point of the line and stays there for a certain time period to heat the material to the intended temperature. When this temperature has been reached, the second phase starts with the movement of the torch along the line with a constant velocity. After a certain distance/time period within the movement the temperature field tends to a quasi-

Transverse stress 11

-250

-150

-50

50

150

250

350

0 5 10 15 20

Depth [mm]

Str

ess

[MP

a]

S4RT C3D8TC3D8RT C3D4TC3D10MT C3D20TC3D20RT

Longitudinal stress 22

-100

100

300

500

700

0 5 10 15 20

Depth [mm]

Str

ess

[MP

a]S4RT C3D8T

C3D8RT C3D4T

C3D10MT C3D20T

C3D20RT

64

stationary state. This means that the temperature field stays constant related to a coordinate system that travels with the centre of the flame. Each point within the quasi-stationary state of the plate runs through a temperature cycle that depends on the transversal coordinates only and not on the longitudinal location.

For a plate t = 20 mm with a torch FB-A 9 the temperature plot is given in Figure 3-6 at a time index of t = 120 s. Generally the location of the maximum temperature is behind the position of the flame centre.

Figure 3-6: Temperatures during the heating process (t = 120 s)

Figure 3-7: Temperatures during the heating process (t = 120 s)

Within the heated area the plastic straining takes place due to the obstruction of the thermal strains by the colder regions in the vicinity. From the stress plots of the simulation it can be seen that in a line heat process this mechanism is three-dimensional (Figure 3-8, Figure 3-9).

Figure 3-8: Stresses in transversal direction during the heating process (t = 120 s)

Figure 3-9: Stresses in longitudinal direction during the heating process (t = 120 s)

In front of the flame an area of tensile stresses forms in transversal direction. But also in longitudinal direction tensile stresses arise near the heated part of the plate. In a section plot (Figure 3-10) it can be seen that the area of tensile stresses forms not only in front of the flame but also below the heated zone.

Figure 3-10: stresses in transversal direction during the heating process (t = 120 s)

The longitudinal distribution of these stresses and plastic strains is shown in Figure 3-11 for the centre line on the heated surface. In front of the heated area the transversal stresses are positive. As soon as the heating by the flame starts the stresses turn into compression. After exceedance of a temperature of about 150° plastic strains start to develop until reaching the maximum temperature of 708° C. Behind this point the level of the plastic strains stays nearly constant.

65

Figure 3-11: Stresses and plastic strains in transversal direction along the centre line during the heating

process (t = 120 s)

Figure 3-12: Development of plastic strains in during the heating process

The plastic strains develop mainly (about 90%) directly under the flame. The remaining part of the plastic strains arises during the cooling phase where the plastic strain curve converges to the final value of the plastic strains (Figure 3-12).

For the tests with short holding time this mechanism is similar, but the plastic strains reach different quantifications. With increasing heat input per unit length (decreasing velocity of the torch) the values for the maximum temperature and maximum plastic strain increase. The recalculation of the small scale tests showed a good agreement with the experimental results. The maximum deviation was less than 5%. Based on the successful recalculation of the simulations have been extended to other nozzles and other plate thicknesses. The results of these parametric variations are shown in Figure 3-13 to Figure 3-16.

As in the small scale tests the maximum temperature is dependent on the heat input per unit length and the thickness of the plate which can be seen from Figure 3-13. Obviously the size and heat flux distribution of the nozzle has also a large influence on the maximum surface temperature. Concentrated heat flux distributions of the single orifice nozzles lead to higher temperatures at lower heat input values than distributed heat fluxes of multi-flame rosebud nozzles. As a consequence, with multi-flame rosebud nozzles higher heat input per unit length can be applied than with single orifice nozzles and, therefore, the resulting rotation angles at the line heat can be increased, what can be seen from Figure 3-14. The direct link between the rotation angles and the heat input per unit length can be seen in Figure 3-15.

Figure 3-13: maximum temperature vs. heat input per unit length

Figure 3-14: rotation angle vs. maximum temperature

To achieve comparability between the different plate thicknesses the heat input per unit length qs is divided by the plate thickness t while the rotation angle is multiplied by t. The result of this transformation is given in Figure 3-16 for the plates with steel grade S355.

66

Figure 3-15: rotation angle vs. heat input per unit length

Figure 3-16: thickness x rotation angle vs. heat input per unit length and thickness

A clear differentiation between the concentrated and the distributed heat flux is shown. Applying the transformation on the results of the small scale tests resp. the numerical simulations with different steel grades the influence of steel grade becomes obvious. For S235 to S460 the interrelation is affine. By dividing the values by the slope m of the regression line, which can be expressed by

4,58 ∙ 10 4,76 ∙ 10 ∙ , fy in [MPa] (3.7)

the influence of the yield strength can be eliminated (see Figure 3-17). The steels S690 and S890 only fit for lower values of the heat input per unit length and thickness up to 200 J/mm² to a linear line. Up to this value a slope of 0,0007 rad mm³/J fits best. All values of m are given in Table 3-3.

Table 3-3 : Slope m for different steel grades (average values)

Steel grade S235 S355 S460 S690 S890

Slope [rad mm³/J] 0,0033 0,0028 0,0022 0,0007 0,0007

The result of this process is given in Figure 3-18 for multi-flame rosebud nozzles with distributed heat flux.

Figure 3-17: thickness x rotation angle vs. heat input per unit length and thickness for different steel grades

Figure 3-18: thickness x rotation angle / m vs. heat input per unit length and thickness

By application of the graph in Figure 3-18 the rotation angle of a heat line can be predicted in dependency of the heat input per unit line, the thickness and the steel grade. For single flame nozzle Figure 3-16 can be transformed into other material grades by multiplying the ordinate with the relation mfy / mS355 given in Table 3-3.

3.2.3 Numerical simulations of T elements

The trials on T-elements (see chapter 2.2.5.2) were numerically simulated and a parametric study was performed in order to evaluate the sensitivity of flame straightening of T element to the burner power, to the heating time and to the oscillation frequency of the burner. The T element was modelled by shell

67

elements. The heat flux was defined by the radiation of a rigid body. The Gaussian distribution of heat flux was considered. The nominal radius of heat source was equal to 41,1 mm, which corresponds to the nozzle no°8 which was used in the trials. The material data were obtained from standard tensile tests (at ambient temp). For higher temperatures the values were obtained in relation to the previously used material model. The initial geometry of T element was perfect and no residual stresses were considered. The FE coupled thermo-mechanical transient analyses were performed for two different heating patterns, for 10 and 20 mm thick plates and for two different material grades.

The first set of FEA was performed on element T5 (material S460, plate thickness 10 mm). The heating pattern with two 30 mm wide (A – see Figure 3-19) heating bands was applied to T5. The positions of the bands were ± 40 mm from the centre (B– see Figure 3-19). In order to shorten computational time, two sources of heat flux were applied simultaneously, as presented in Figure 3-20.

Figure 3-19: Definition of heating areas

Figure 3-20: Surface temperature at 45 s (element T5)

Figure 3-21 displays the results of the parametric study for element T5. The varied parameters were:

- heating time: time of the heating source to travel from the top to the bottom edge of the element;

- power of burner: by varying the temperature of the heating source, the effective heating flux varies.

The curves with symbols in Figure 3-21a present the final displacement (E – see Figure 3-19), while the bold lines present the maximum surface temperature (y-axis on the right) in relation to the power of burner. Each curve presents heating time (yellow 60s, blue 120s, magenta 240s and red 480s). The final displacement rapidly increases if the burner power increases. If heating times are longer, the final displacement reaches an upper limit at burner powers that are commonly used in the flame straightening processes (burner nozzle n°6 and larger). This happens due to the plate overheating. In Figure 3-21b the displacements and temperatures are plotted in relation to the energy input. Higher energy input does not always results in higher final displacement. The oscillation of the burner (period time 1s, 3s and no oscillation) within the heating band has no significant effect on the final displacement for burner powers 4 to 14 kW.

68

a) b)

Figure 3-21: The effect of heating time, power of the burner on the displacement and surface temperature for element T5

The comparison of experimental results, shown in Figure 3-22, and results obtained by FEA is presented in Figure 3-23. The curve with symbols in Figure 3-23 presents the final displacement in relation to the burner power for 120s heating time, while the remaining lines present the maximum temperatures in different depths of the plate.

Figure 3-22: Basic input data and experimental results for element T5

Element T5 was flame straightened in two cycles. The heating times were between 102 and 135 s for burner path on one side of the element. The maximum measured temperature in the steel plate 0,7 mm below the flame (thermocouple T5 in Figure 3-22) was approximately 600°C. The final displacements on each side of T element in the centre were 4,2 and 5,7 mm in the first cycle and 4,1 and 4,2 in the second cycle. If a horizontal line is drawn in Figure 3-23 from maximum measured temperature equalling 600°C to the temperature line on the appropriate depth (see interpolated temp. line in Figure 3-23), and if then vertical line is drawn from the point on temperature line to the curves with symbols, the effective burner power and the final displacement as a result of numerical simulation may be read. It follows that the measured temperature of 600°C corresponds to numerically obtained final displacement 4,2 mm and to the power of the burner 7900 W. This agrees with experimental results very well. The effective power of the burner that was used in the trials was estimated to 7800 W.

0.0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

8.0

9.0

10.0

0 2000 4000 6000 8000 10000 12000 14000 16000Power of burner [W]

Dis

plac

emen

t [m

m]

0

200

400

600

800

1000

1200

1400

1600

Tm

ax (

Top

sur

face

) [°

C]

Disp. 60s Disp. 120sDisp. 240s Disp. 480sTmax top 60s Tmax top 120sTmax top 240s Tmax top 480s

0.0

1.0

2.0

3.0

4.0

5.0

6.0

7.0

8.0

9.0

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0 500 1000 1500 2000 2500 3000 3500Energy input [kJ]

Dis

plac

emen

t [m

m]

0

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800

1000

1200

1400

1600

Tem

pera

ture

[°C

]

Disp. 60s Disp. 120sDisp. 240s Disp. 480sTmax top 60s Tmax top 120sTmax top 240s Tmax top 480s

Width of the heating band [mm] A [mm] 30 30 30Position relative to centre of the web B [mm] -40 -40 40Time of heating[s] C [s] 135 121 102Displacement per cycle [mm] E [mm] 4.0 4.0 3.6

E [mm] 4.2 4.1 4.2

DescriptionCycle

4.55.7

3040113

1 2Specimen

T5

69

Figure 3-23: Comparison of experimental and FEA results for element T5

A series of FEAs was performed on element T2 (material S235, plate thickness 10 mm). The heating pattern with a single 120 mm wide heating band was applied to the element. A single heating source oscillated within the heating band with the period of 5s. The curve with symbols in Figure 3-24 presents the results for heating time equalling 240 s and for different power of the burner. The figure also presents temperatures over the plate thickness. The final displacements measured in the trial were 3,6 and 4,1 mm on each side of the flange. The measured displacement corresponds well to the FEA results if the estimated heat input is the starting point for the comparison (see red lines with arrows in Figure 3-24). If the starting point for the comparison of FEA and experimental results is the temperature (see green lines with arrows in Figure 3-24), the final displacement should be approximately 5 mm.

Figure 3-24: Comparison of experimental and FEA results for element T2

Another series of FEA was performed for element T6 that was assembled with 20 mm thick plates. The results of the experiment are presented in Figure 3-25. The final displacements achieved in the experiment in the first cycle were 2,3 to 3,1 mm at heating times equalling 130 and 123 s, while the displacements in the next two cycles were on average smaller despite longer heating times. This does not correspond to the FEA results. In Figure 3-26 the curves with symbols present the displacement of flange for heating times equalling 120 and 240 s. The reason for this disagreement is being studied, but one of possible answers is that the second and third cycles of the trial started when the initial temperature was 100 to 120 °C. The temperature gradient over the plate thickness is important for effective flame straightening of flanges of T profiles.

0.0

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plac

emen

t [m

m]

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Tm

ax (

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sur

face

) [°

C]

Disp. 120sTmax Top surfaceTmax 8.33mmTmax 6.67mmTmax 5mmTmax 3.33mmTmax 1.67mmTmax Bottom surface

interpolated temp. line

experimental Tmax = 600°C

0.0

1.0

2.0

3.0

4.0

5.0

6.0

6000 7000 8000 9000 10000 11000 12000

Power of burner [W]

Dis

plac

emen

t [m

m]

0

100

200

300

400

500

600

700

800

T [

°C]

heating time = 240s

Temp Bottom

Temp 1.67mm

Temp 3.3mm

Temp 5mm

Temp 6.67mm

Temp 8.33mm

Temp Top

Measured temp. in the experiment (8.9 mm)

Estimated heat input

interpolated temp. line

70

Figure 3-25: The basic input data and the final displacement for element T6 for the experiment

Figure 3-26: The effect of heating time, power of the burner on the displacement and surface temperature for element T6

3.3 Bar shaped structures

Structures composed of beams have been investigated by the means of finite element simulation to identify the mechanism of the bending processes performed in the experimental part of the project.

3.3.1 Transient, fully coupled thermo-mechanical numerical simulations of beams

A series of numerical simulations was performed in order to evaluate the influence of initial residual stress, the magnitude of external load, the size of heating pattern and the heating sequence on the final displacement. The simulations are based on the trial 4.1-13, 6 m long IPE450 beam, to which V heats were applied to derive a bending about the strong axis. The positions and size of V heats are shown in Figure 3-27. The numerical model of burners is presented in Figure 3-28.

Figure 3-27: The positions of V heats on the beam Figure 3-28: The numerical model of burners

Width of the heating band [mm] A [mm] 35 35 35 35 35Position relative to centre of the web B [mm] -43 -43 43 -43 43Time of heating[s] C [s] 130 245 185 192 200Displacement per cycle [mm] E [mm] 2.3 1.6 2.5 1.0 0.5

E [mm] 3.1 1.7 2.1 1.8 1.4

Despription

2.32.7

Cycle

3543123

1 2 3Specimen

T6

0.0

1.0

2.0

3.0

4.0

5.0

1000 3000 5000 7000 9000 11000 13000 15000Power of burner [W]

Dis

plac

emen

t [m

m]

0

200

400

600

800

1000

1200

Tm

ax (

Top

sur

face

) [°

C]

Displacement 120s

Displacement 240s

Tmax top 120s

Tmax top 240s

71

Figure 3-29: The temperature dependent stress-strain curves as input data for numerical simulations

In all the simulations the mean material values for S235 were assigned to the material model. The temperature stress-stain relationship is presented in Figure 3-29.

The residual stresses were accounted in the numerical simulations only for the longitudinal direction of the beam axis (Figure 3-30). Its magnitude was determined according to the recommendations in [2].

Figure 3-30: The initial residual stress in the flange and web of IPE450 according to the literature [2] (ECCS) and as an input data in the numerical analyses (Abaqus)

The input data for all numerical simulations is presented in Table 3-4. The power of burners are lower than the effective burner power used in the trial because the oscillation of the burner is taken into account by larger burner area that covers whole width of V heat (see Figure 3-28). The heat flux (in units J/(s m2)) is equal in all analyses.

Figure 3-31 shows the temperature in points A and B through the heating phase (to 210 s) for the analysis with 80 mm wide V heat. The four passes of burner over the flange are followed by heating of the web from one and the other side. Uniform temperature over the plate thickness reached in a short time period (approximately 15 s). Moreover, the temperature on the heating surface is much larger than the uniform temperature. The temperature diagram for wider V heats (120, 160 mm) and the same heating sequence have no more than 50°C higher local maximums.

0.0

100.0

200.0

300.0

400.0

500.0

600.0

0.000 0.050 0.100 0.150 0.200 0.250True plastic strain

Tru

e st

ress

[MP

a]

0 80 180 280 380 480

580 780 680 880 980

Flange

-75

0

75

150

-95 -45 5 55

coordinate [mm]

Str

ess

[MP

a]

Abaqus

ECCS

Web

-210

-140

-70

0

70

140

-210 -140 -70 0 70 140 210

coordinate [mm]

Str

es

s [

MP

a]

Abaqus

ECCS

72

Figure 3-31: The temperature in point A and B only for the heating phase for analyses S1, S1RS, S5 and S5RS

Figure 3-32: Vertical displacement (the influence of the initial residual stress and heating sequence)

Figure 3-33: The deflection of the beam in heating and cooling phase and the final deflection (the influence of the initial residual stress, width of V heat and external load)

0

200

400

600

800

1000

1200

0.00 52.50 105.00 157.50 210.00

Time [s]

Tem

per

atu

re

Point A upper surf.Point A 3.65 mm below upper surfPoint A bottom surf.Point B front surf.Point B back surf.

-10

-7.5

-5

-2.5

0

0 1000 2000 3000 4000 5000 6000

Longitudinal axis [mm]

Ver

tica

l dis

pla

cem

ent

[mm

]

S1 S1 RSS2 S2 RSS3 S4

S1

S1

RS

S5

S5

RS

S1

RS

F80

S1

RS

F10

0

S1

RS

b12

0

S1

RS

b12

0 F

80

S1

RS

b12

0 F

100

S1

RS

b12

0 F

120

S1

RS

b16

0

S1

RS

b16

0 F

80

S1

RS

b16

0 F

100

-22

-17

-12

-7

-2

3

8

Name of the analysis

Dis

pla

cem

ent

[mm

]

heating phasecooling phasefinal displacement

73

Figure 3-34: Final vertical displacements in dependence to the external force for different widths of V heat

a) S1_RS b) S1_RS_b160_F100

Figure 3-35: Equivalent plastic strain after unloading (Deformation Scale Factor = 10 ) in the mid-section point

Figure 3-32 illustrates the vertical displacement along the beam longitudinal axis. The heating sequence has an influence on the displacement (analyses S1 to S4 – see also Table 3-4). If the residual stresses are considered, the influence of the heating sequence becomes less important. The introduction of the residual stress in the numerical model results in lower final displacement.

Figure 3-33 presents final displacements (in the middle of the beam) as well as the displacements at the end of heating and cooling phase. The deflection of the beam in the cooling phase is independent, while the lift of the beam in the heating phase is dependant of the external force and the initial residual stress. In two cases (wide V heats and high external load) the deflection occurred in the heating phase due to the formation of a plastic hinge in the V heats. Figure 3-34 presents the displacements in dependence of the normalized plastic resistance for different widths of V heat. The external force increases the displacement linearly until the formation of plastic hinge. From that point forward, the displacement increases exponentially. In Figure 3-35 the equivalent plastic strain (PEEQ)is shown for two simulations. Although the final displacement was only 4,9 mm in Figure 3-35a the PEEQ locally exceeds 5%. The formation of plastic hinge in Figure 3-35b causes flange and web instabilities with large PEEQ.

-25

-20

-15

-10

-5

0

0.00 0.05 0.10 0.15 0.20 0.25 0.30

M/Mpl (fy = 329 MPa)

Fin

al d

isp

lac

em

en

t [m

m]

80

120

160

Width of V heat in mm

74

Table 3-4: Input data and final displacement in numerical simulations of IPE450

Fin

al

disp

lace

men

t [m

m]

-7.4

-4.9

-8.3

-5.2

-6.4

-7.4

-4.5

-2.3

-5.4

-7.2

-6.2

-3.8

-8.3

-10.

6

-15.

7

-9.1

-3.8

-11.

9

-13.

4

-21.

6

1 fl

ange

4 p

asse

s in

105

s, f

ollo

wed

by

heat

ing

of th

e on

e si

de o

f w

eb in

52.

5s, a

nd th

e ot

her

side

of

the

web

in 5

2.5s

1a

fl

ange

4 p

asse

s in

105

s, f

ollo

wed

by

heat

ing

of th

e on

e si

de o

f w

eb in

52.

5s, a

nd th

e ot

her

side

of

the

web

in 5

7s

2 tw

o pa

sses

ove

r th

e fl

ange

in 5

2.5

s, f

ollo

wed

by

sim

ulta

neou

s he

atin

g of

the

flan

ge in

two

pass

es a

nd b

oth

side

s of

the

web

(si

mul

tane

ousl

y)

3 tw

o pa

sses

ove

r th

e fl

ange

and

a s

ingl

e pa

ss o

ver

one

side

of

the

web

in 5

2.5

s, p

roce

dure

rep

eate

d in

the

follo

win

g 52

.5 s

4

flan

ge 4

pas

ses

in 1

05 s

, fol

low

ed b

y 12

0 s

of n

o he

atin

g (c

ondu

ctio

n in

the

beam

), f

ollo

wed

by

heat

ing

both

sid

es o

f th

e w

eb s

imul

tane

ousl

y in

52.

5s

Inpu

t en

ergy

per

V

hea

t [k

J]

web

399

399

399

399

399

399

399

399

399

399

399

598

399

598

598

598

797

797

797

797

flan

ge

399

399

399

399

399

399

399

399

399

399

399

598

399

598

598

598

797

797

797

797

Hea

ting

tim

e [s

]

web

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

flan

ge

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

105

Pow

er [

W]

of

burn

ers w

eb

(eac

h)

3796

3796

3796

3796

3796

3796

3796

3796

3796

3796

3796

5694

3796

5694

5694

5694

7592

7592

7592

7592

flan

ge

3796

3796

3796

3796

3796

3796

3796

3796

3796

3796

3796

5694

3796

5694

5694

5694

7592

7592

7592

7592

V h

eat

wid

th

[mm

]

80

80

80

80

80

80

80

80

80

80

80

120

80

120

120

120

160

160

160

160

Sequ

ence

of

hea

ting

1 1 2 2 3 4 1 1 1 1 1 1 1 1 1 1 1 1 1 1

Ext

erna

l fo

rce

P

[kN

]

68

68

68

68

68

68

0 0 80

120

100 0 68

100

120

80

0 68

80

100

Res

idua

l st

ress

no

yes

no

yes

no

no

no

yes

yes

yes

yes

yes

yes

yes

yes

yes

yes

yes

yes

yes

Ana

lysi

s na

me

S1

S1R

S

S2

S2 R

S

S3

S4

S5

S5 R

S

S1R

S_F8

0

S1R

S_F1

20

S1R

S_F1

00

S1R

S_b1

20_F

0

S1R

S_b1

20

S1R

S_b1

20_F

100

S1R

S_b1

20_F

120

S1R

S_b1

20_F

80

S1R

S_b1

60_F

0

S1R

S_b1

60

S1R

S_b1

60_F

80

S1R

S_b1

60_F

100

No.

1 2 3 4 5 6 7 8 9 10

11

12

13

14

15

16

17

18

19

20

3.3.2 Steady-state, fully coupled thermo-mechanical numerical simulations of beams

In order to study the phenomena, the focus is set on heating of a single V located in the middle of the beam. The numerical and material model of the beam is the same as in previous Section 3.3.1. The numerical simulation was performed in four steps. In the first step (t = 1) the external force was applied to the beam. This was followed by coupled thermo-mechanic steady state step (heating phase), where the temperature in the heating zone was prescribed. In the third step the temperature of the heating zone was prescribed back to the initial value (cooling phase). The temperature outside the heating zone was equal to the initial temperature throughout the heating and cooling phase. In the last step the beam was unloaded.

75

Figure 3-36 shows the development of the maximum displacement of the beam during the four steps. In the heating phase the beam is lifting due to expansion of the flange. After reaching the first critical temperature Tcrit,1 the lifting of the beam stops. At this point the compression stress in the heated zone has reached the proportional limit fp,Tcrit,1 and the moment resistance of the V heat cross-section is equal to the moment cause by external load. The equilibrium is assured by the tension stress in the unheated part of the cross-section (see Figure 3-37). When the temperature increases the material point reaches (drops to) different stress-strain curves at the same stress, but higher plastic strain. The stress distribution in the cross section cannot longer change, when T > Tcrit,1, because the external force is constant. At the second critical temperature, the heated part of the cross section reaches yield stress fy,Tcrit,2. At fy,Tcrit,2 the plastic strain in the heated part becomes around 2%. If the temperature in the heated zone further increases, the cross-section in the heated zone cannot longer bear the stress and equilibrium is assured more or less only by the unheated part of the cross-section until it fully yields (see Figure 3-37). This phase results in large displacement of the beam and buckling of the web and the flange. In the cooling phase the heated part contracts due to temperature strain. The plastic strain is irreversible and results in the deflection of the beam.

Figure 3-36: The development of the vertical displacement in coupled steady-state thermomechanical

analysis

Figure 3-37: Normal stress component in the cross-section at different temperatures of V heat (height of

V heat form z = –225 to z = 75)

3.3.3 Methodology for 3D-modeling of the flame bending process of beams

3.3.3.1 Procedure

Simulating a complete heat straightening process of a large beam profile with several heated areas, it can be very complex from a simulation point of view. The time and computational requirements for these simulations make them unviable. However, the methodology outlined here, developed for use with ANSYS finite elements code, makes this process affordable.

A typical sequence for a heat straightening in a beam consists of applying the flame in several areas with different heating patterns, such as V in the web and transversal bands in the flange. Bearing in mind this typical working procedure, the main foundation of this method consists of simplifying the entire beam into a set of short beams, one for every heated area, considering these as independent from each other. Figure 3-38 shows how the entire beam is discretized into a number of those short beams.

-20

-18

-16

-14

-12

-10

-8

-6

-4

-2

0

0 1 2 3 4Pseudo time

Dis

pla

ce

me

nt

[mm

]

Ap

ply

ing

ext

ern

al f

orc

e

Hea

tin

g p

has

et

= 1

equ

als

to 2

0 °C

t =

2 e

qual

s to

800

°C

Co

oli

ng

ph

ase

t =

2 e

qual

s to

800

°C

t =

3 e

qual

s to

0°C

Un

load

ing

th

e b

eam

Tcrit,1

(c) -> (e) Formation of plastic hinge -> kinematic instability

Tcrit,2

(a)

(b)

(c)

(d)

(e)

-225

-75

75

225

-330 -210 -90 30 150 270

Stress [MPa]

z [m

m]

Tcrit,1 = 612 °CTcrit,2 = 670 °CT = 749 °CT = 799 °C

76

Figure 3-38: Simplification of an entire beam into several short beams

This method involves other reasonable simplifications including the following boundary conditions: - Convection was applied on every area with a film coefficient of 10 W/m2K and bulk

temperature was set to 25 ºC. - The edges of the short beam must preserve their original shape to guarantee the mechanical

independency from the heated area. Thus, the finite element simulations are performed in two steps. First of all, the short beam is subjected to a thermal simulation and then, with the data provided by the first simulation, the mechanical simulation starts. Bearing in mind this separation of simulations, for an entire beam, only one thermal simulation must be developed because all the heated spots are usually equivalents. The mechanical process must be developed for each heated area because the bending moment to introduce in each case is different because it depends on the position the short beam take up in the entire beam.

The simplification of the entire beam separated into several short beams relies on the thermal and mechanical independency between each short beam and the next one. The thermal independency has been checked with experimental methods and then the thermal and the mechanical independency have been verified with the data provided by the simulation.

The experimental tests developed to check the thermal independency consisted in subjecting a steel plate to a flame applied through a band (see Figure 3-39) and measuring the temperature. The temperature was measured near the heated band, 150 mm far from it and on the surface of the steel and it was not higher than 60 ºC. At this temperature the steel properties are not affected at all so the thermal independency for longer distances than 150 mm from the flame is guaranteed.

Figure 3-39: Experimental verification for the thermal independency over a steel plate.

Once this simplification is accepted, the simulation of the entire beam is carried out in several simple simulations, one for each short beam, and it is affordable.

After the entire simulation has been completed, the data provided by it allows the final deflection in the beam by the integration of all of the partial deflections of each beam to be estimated. The partial

F

77

deflection of each beam can be calculated as a function of the position it takes up in the entire beam and the rotation angle it has after the simulation, as can be observed in Figure 3-40.

Figure 3-40: Sketch of the estimation of the rotation angle in the short beam.

Figure 3-41: Final deflection sketch

Once the partial deflection for each short beam is known, the final deflection can be easily calculated by applying the equations of trigonometry considering the spot where the flame is applied as a plastic hinge (Figure 3-41).

To simulate a complete beam, a mechanical simulation must be performed for every heated area with its equivalent bending moment value. However, as it is visible in Figure 3-42 where it can be observed the relationship between the applied bending moment and the corresponding rotation angle for three different temperature levels, the values for bending moments and temperature normally applied fall into a straight line or follow a simple-fitting trend, so an entire beam can be simulated only with the values of two or three heated areas. The rest of rotation angle values can be obtained simply by interpolating in the graph provided by these few points.

Figure 3-42: HEA300 Bending moment-Rotation graph

3.3.3.2 Experimental calibration of the models

To evaluate the precision of this methodology many laboratory tests were carried out. In this report the results for three of these tests are summarised and compared with the total deflection values obtained by ANSYS using the proposed methodology. Each of these beams was subjected to different additional

HEA 300

0.0E+00

1.0E-03

2.0E-03

3.0E-03

4.0E-03

5.0E-03

6.0E-03

7.0E-03

0 20000 40000 60000 80000 100000 120000 140000

Bending Moment N*m

Ro

tati

on

(ra

dia

ns)

HEA 300 1000ºC

HEA 300 850ºC

HEA 300 760ºC

0.25Mp

0.125Mp

R =R(Mb)

∆θ

78

loads in these laboratory tests. These loads had different values depending on the steel grade of the beam (S235J0, S355JR and S460ML) and on the profile (HEA 300, IPE 450 and HEM 340). The heating patterns were “V” shape in the web and a transversal band in the flange (see Figure 3-43). The beam was discretised into a number of short beams, one for each heated area. The short beams were 50 cm long, 25 cm either side of the centre, long enough to ensure that the edges are not significantly affected thermally or mechanically by the flame.

Figure 3-43: Heating patterns, discretised area

The first test was a HEA 300 profile, S355J2 steel grade, consisting in two steps (Figure 3-44), the first one (red) with an additional load of 70 kN and the second step without any additional load. The second test was a HEA300 profile, S460ML steel grade, also consisting in two steps (Figure 3-45), the first with an additional load of 40 kN and the second without any additional load. The third test (Figure 3-46) was an IPE 450 S235JR profile and consisted of three steps, the first without any additional load (vees 1, 2 and 3), the second with an additional load of 75 kN (vees 4, 5, 6 and 7) and the third step with an additional load of 67 kN (vees 8, 9, 10 and 11). Table 3-5, Table 3-6 and Table 3-7 summarised the results of all these tests and their corresponding numerical simulations.

Test 1:

Figure 3-44: HEA 300 S355 test setup

Step 1 (deflection mm) Step 2 (deflection mm) Final Results (deflection mm) ANSYS Real Test ANSYS Real Test ANSYS Real Test

26.6 29.8 8.5 11.6 35.0 41.4

Table 3-5: Summary of the results.

Test 2:

Figure 3-45: HEA 300 S460 test setup

79

Step 1 (deflection mm) Step 2 (deflection mm) Final Results (deflection mm) ANSYS Real Test ANSYS Real Test ANSYS Real Test

17.9 31.0 9.5 16.7 27.3 47.7

Table 3-6: Summary of the results.

Test 3:

Figure 3-46: IPE 450 S460ML test setup

Step 1 (deflection mm)

Step 2 (deflection mm)

Step 3 (deflection mm)

Final Results (deflection mm)

ANSYS Real Test

ANSYS Real Test

ANSYS Real Test

ANSYS Real Test

2.7 1.3 8.2 10.0 8.0 9.6 18.8 21.1

Table 3-7. Summary of the results.

80

4 Analytical investigations

With the results of the experimental investigations and the numerical simulations the mechanisms of the heating patterns have been further developed for bar shaped structures. Two models have been derived which are presented hereafter.

4.1 Presentation of analytical model for bar shaped structures

4.1.1 Model 1: Critical temperature approach

This analytical model was derived for the beam elements, heated with V heats or some other shape of localized heat. The main suppositions of this model are:

- the heating in localized (V shape, vertical band or similar); - the heated part of the beam has uniform temperature field, while the temperature field of the

unheated part is not affected by heating; - the initial stress due to external load or gravity causes permanent deformations; - the neutral axis forms between heated and unheated part of the cross-section.

The result of this model is the critical temperature of the heat that results in significant deformation of the beam at a prescribed level of external load.

Let the cross-section in the V heat be considered. The beam is loaded with external load that causes the bending moment Mvee in the V heat (Figure 4-1). The moment induces linear stress distribution in the cross section with the maximum value of 0 (state (a) in Figure 4-2). Height hv of this cross section is exposed to V heat. The first critical temperature is reached when the moment resistance of the cross section is equal to the bending moment Mvee caused by the external force. At this point the stress at the heated part of the cross section reaches the proportional limit fp,Tcrit,1. The tension in the unheated part equilibrates the compression in the heated part. The assumption is that the unheated part of the cross section has initial temperature of the beam before heating (room temperature). It follows:

,1, crit

veep T

c T

Mf

A c , (4.1)

where

Mvee moment due to external force at x = lv,

Ac area of the cross section in the compression,

cT the distance between centres of gravity of the cross-section in compression and in tension.

The second critical temperature is determined in a similar way. The increase of the temperature from Tcrit,1 to Tcrit,2 does not change the stress distribution in the cross-section, but the there is a significant increase of deformation in the heated part.

,2 ,1, ,crit crit

veey T p T

c T

Mf f

A c (4.2)

81

Figure 4-1: Deformation of the beam due to plastic strain in V heat

Figure 4-2: Simplified stress distribution in the V heat at different temperature levels

The critical temperatures are calculated from the material characteristics. Therefore, the temperature should be written as a function of proportional limit (for Tcrit,1) and at yield stress (for Tcrit,2), respectively, or may be extrapolated from EN 1993-1-2, Table 3.1.

The deformation of the beam may be calculated according to the following method. Let us assume that deformation is concentrated in the V heat, which acts as a nonlinear linear spring, while the beam outside V heat acts as a rigid body. The width of V heat after cooling bv,T may be expressed by the contracting strain T (positive sign stands for compression), as (see Figure 4-1):

, 1v T T vb b (4.3)

The temperatures higher than Tcrit,2 cause plastic hinge in the heat, resulting in large strain (infinite strain in case of perfect plasticity). Therefore, the moment resistance of the beam is restricted to the unheated part of the cross-section (see Figure 4-2 (d), (e)). Moreover, temperature higher than Tcrit,2 would result in uncontrolled sagging of the beam, connected with local and global instabilities. This happens in case when the assumption about uniform temperature field is fulfilled. Due to high conductance, it is difficult to unintentionally achieve the uniform temperature field by flame heating method. When the heating torch passes over a material point near the heated surface it generates high temperature field. The temperatures higher than Tcrit,1 cause irreversible plastic strain. Due to high conductance the temperature quickly drops pushing the material points into the elastic part of the material diagram. Therefore in the end of the heating phase the beam is not significantly deformed. As presented in 3.3.2 the temperatures below Tcrit,1 cause lifting of the beam in the heating phase. In the cooling phase the beam sags due to temperature strain. Therefore it is reasonable to assume that the thermal contraction causes the final displacement of the beam. The strain in the V heat may be written as:

T T T (4.4)

The deformation of the beam may be expressed by rotations in the V heat (see Figure 4-1).

82

1 2 (4.5)

(4.6)

2arctan2

v

v

b

h

(4.7)

Because is very small, the simplification = tan is in order and may be written as:

v

vT

v

Tv

h

b

h

b

2

1arctan2

2arctan2 , (4.8)

The displacement of the beam under the V heat may be expressed by the rotation in V heat:

1 2

v vVee v v

l l l lw w l l

l

(4.9)

If two V heats are positioned symmetrically to the external force and external force is located at l/2, the maximum deflection w is:

Vee vw w l (4.10)

To control the results of this procedure the stiffness of the unheated part is considered. If the stiffness of the heated part is neglected (see Figure 4-2 (e)), the unheated part resist to the external load. The final estimation for the displacement W of the beam is:

min ;Vee UnheatW w w , (4.11)

where wUnheat is the elastic displacement of the beam with only unheated part of the cross-section.

4.1.2 Model 2: Determination of the deformation of bar-shaped elements by flame straightening using analytical prediction means

In the following an analytical method is presented with which the deformations of the flame bending process of bar-shaped elements can be estimated.

4.1.2.1 Considered change in material properties

During the heating process parts of the section of the beam are heated over temperatures of 600° C which leads to significant changes in material properties:

The yield strength fy resp. the proportionality limit fp decreases with increasing temperature T. This decrease is approximated by a hyperbolic tangent function in the following form:

RTyy fT

Tf ,4,282

4,575tanh1493,0

(4.12)

Beside the strength of the material also the Young’s modulus E decreases with temperature which is also given by a hyperbolic tangent function:

RTET

TE

0,329

8,692tanh1498,0

(4.13)

Both functions are derived by non-linear regression from data derived by hot tensile tests and are given in Figure 4-3 and Figure 4-4.

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Figure 4-3: Ratio of Young’s Modulus vs. temperature Figure 4-4: Ratio of yield strength vs. temperature

Even if the thermal expansion coefficient depends on the temperature it is assumed to be constant for the further investigations. Other changes (e.g. material density ) are not considered as there effect on the deformation is negligible.

4.1.2.2 Presentation of analytical model 2

To determine the resulting deformations from the heating process the knowledge of the distribution and size of the occurring plastic strains is required. In the following we assume a bar-shaped element with the coordinate system and internal forces given in the Figure 4-5.

Figure 4-5: coordinate system and internal forces Figure 4-6: Thermal strains

In the heated part of the beam section thermal strains occur which depend on the thermal expansion coefficient and the difference between the actual and the ambient temperature:

0TTthth (4.14)

Due to

- the stiffness of the remaining cooler section of the heating pattern

- the stiffness of the overall system

- the internal forces in the section of the heating pattern arising from external loads on the element

compressive stresses arise in the heated part. If these compressive stresses exceed the yield strength relevant for the actual temperature plastic shortening occurs. To determine these compressive plastic strains the strain state at different time stages is investigated and the resulting plastic strains are estimated.

Principally the strain state consists of two different parts:

0,0

0,2

0,4

0,6

0,8

1,0

1,2

0 200 400 600 800 1000

Rat

io E

/ER

T

Temperature T, ° C

data

tanh

0,0

0,2

0,4

0,6

0,8

1,0

1,2

0 200 400 600 800 1000

Rat

io f

y/f

y,R

T

Temperature T, ° C

data

tanh

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- thermal strains th due to the heating process

- mechanical strains m

- due to the reaction of the section to the thermal strains,

- due to external forces (normal forces, bending moment, etc.),

- due to the reaction of the overall system,

where for the mechanical strains the hypothesis of Bernoulli is assumed. Therefore the mechanical strains form a plane which can be defined by three sampling points m1, m2 and m3 which are so far unknown.

All strain parts are superimposed in the considered section to a total strain distribution:

partheatedtheinzyzy thmtot ,, partheatednontheinzyzy mtot ,,

The total strains are linked to the stresses by constitutive equations, e.g.

TE

TfifTf

TE

TfifzyTE

zyy

ytotytot

yytottot

sgn

,

, (4.15)

for linear-elastic ideal-plastic material behaviour.

Figure 4-7: Stress distribution

The stress distribution has to fulfil equilibrium with the internal forces given by the equations

NdA (4.16)

yMdAz (4.17)

zMdAy (4.18)

Using equations (4.16), (4.17) and (4.18) the unknown sampling points m1, m2 and m3 can be derived. Having the three sampling points the total strain distribution can be calculated and, if the total strain exceeds the yield strain y, the plastic strain increment is given by:

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ytot

ytotytot

pl if

ifzyzy

0

,, (4.19)

After cooling the arisen plastic strains pl affect a corresponding mechanical strain m. The superposition of both strains results in a stress distribution which has to fulfil all conditions for residual stress distribution which means the considered section has to be in equilibrium:

0 dA (4.20)

0 dAz (4.21)

0 dAy (4.22)

Using the boundary values of the resulting mechanical strain distribution the curvature can be calculated using the height h of the beam

hlowermupperm

y,,

(4.23)

brightmleftm

z,,

(4.24)

Using this procedure to determine the curvature for several sections along the heating pattern the resulting rotation can be calculated by integration of the curvature y or z by x using the principle of work

dxMxyy with 1M (4.25)

dxMxzz with 1M (4.26)

A fibre model of different sections including the described procedure has been established in an Excel-Worksheet and calibrated by the results of the large scale trials. Figure 4-9 and Figure 4-8 show the comparison of the deformations gained by the analytical approach and the experimental results of the large scale trial using constant temperature fields selected according to those measured during the tests.

In general the analytical results show a good agreement with the experimental results. The procedure helps to further understand the processes and the results arising from the flame bending process. For a vee heat process on an IPE 450 profile, steel grade S235 with a vee heat about the strong axis (hvee = 300 mm, bvee = 80 mm) the strain and stress distributions in the center section of the vee heat are shown for the time step of the maximum temperature in Figure 4-10.

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Test 4.1-13 Tmax = 890 ° C fy = 338 MPa IPE 450 S235

Figure 4-8: Comparison of analytical and experimental results for tests 4.1-13

Test 4.1-14

Tmax = 1000 ° C fy = 338 MPa IPE 450 S235

Test 4.1-17 Tmax = 950 ° C fy = 484 MPa IPE 450 S460

Figure 4-9: Comparison of analytical and experimental results for tests 4.1-14 and 4.1-17

0,0

5,0

10,0

15,0

20,0

25,0

0 1000 2000 3000 4000 5000 6000

vert

ica

l dis

pla

cme

nt, m

m

Longitudinal axis, mm

analytical approach

after flame straightening process

0,0

5,0

10,0

15,0

20,0

25,0

30,0

0 1000 2000 3000 4000 5000 6000

vert

ical d

isp

lace

me

nt, m

m

longitudinal axis, mm

analytical approach

after flame straightening process

0,0

5,0

10,0

15,0

20,0

25,0

0 1000 2000 3000 4000 5000 6000

vert

ical d

isp

lace

me

nt, m

m

longitudinal axis, mm

analytical approach

after flame straightening process

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Thermal and mechanical strains Stresses Plastic strains

M = 50 kNm

M = 95 kNm

Figure 4-10: strain and stress distributions (IPE 450 profile, S235, vee heat: 300 mm/80 mm)

From this figures it can be seen that due to the increasing bending moment the mechanical strain distribution changes from positive to negative values and due to this the stress distribution in the not heated part alters completely. The resulting plastic strains increase from about 0,5% to at least 2%. By further varying of the external moment the formation of the plastic hinge can be approximated (see also numerical simulation) and it is possible to derive a moment-rotation curves dependent on different parameters (see also Figure A2-7.27 and Figure A2-7.28 in chapter 7.4 of the guideline in Appendix 2).

For a vee heat about the strong axis of I section the curvature depends on the height of the heating pattern at the location. The curvature is affine to the first moment of area of the heated part of the section (Figure 4-11). Instead of a vee heat for a longitudinal line heat at the centre of the flange of a profile the curvature is continuous and increases from the supports to the middle of the beam (see Figure 4-12).

Figure 4-11: Distribution of in a vee heat Figure 4-12: Distribution of for a longitudinal line

heat at the centre of the flange

For the calibration experimental results have been used with detailed information on geometry and sequence of the heating pattern and the temperature distributions (see Figure 4-13 and Figure 4-14). For the estimation of the deformation in most cases the exact temperature field is unknown and can only be predicted by numerical analysis.

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Figure 4-13: Temperature distribution on the surface of the flange

Figure 4-14: Temperature distribution on the surface of the web

The accuracy of estimation by the described procedure depends mainly on the knowledge of the temperature distribution within the heating pattern at different time steps. Special attention should be drawn towards the sequence of heating e.g. the sequence of the heating of the web and the flange within a vee heat.

4.2 Comparison of analytical and numerical prediction means

As shown before both analytical models and the numerical model have been calibrated using the experimental results gained by the large scale tests. Nevertheless the analytical models are based on several simplifications and assumptions that may lead to a deviation from the real deformation. To give a comparison of the three methods a profile IPE 450 in steel grade S235 has been selected as reference case. With this profile a bending about the strong axis shall be investigated with a vee heat with 300 mm of height and 80 mm of width. An external force is supposed to induce a bending moment about the strong axis in different quantifications from 0 to 25% of the plastic bending moment of the section. With model 1 the critical temperatures Tcrit,2 have been calculated. These values are given in the graph the vertical dotted lines for the different bending moments. With model 2 the blue curves have been derived under the assumption of linear elastic-ideal plastic material behaviour and a constant temperature in the vee heat. By the numerical model the red curves have been derived by the method described in 3.3.3. All results are given in Figure 4-15

By the numerical models a continuous relationship between the rotation and the temperature has been calculated. With increasing temperature the rotation increases also. Compared to this the results of the analytical models are discontinuous. Up to the critical temperature Tcrit,2 of model 1 the result of model 2 is line with a very small slope. Above the critical temperature, the rotations start to increase with increasing temperatures in a slope larger than that of the numerical results. In comparison the size of the rotations for both numerical and analytical model fit well.

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Figure 4-15: Comparison of analytical and numerical rotations in relation to the temperature in the vee heat

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5 Conclusions

Within the project by experimental, numerical and analytical investigations the flame straightening process of different structural elements has been examined thoroughly and a comprehensive inside look of the process has been worked out. In the following the main conclusions of these investigations are given divided into the experimental, numerical and analytical part.

5.1 Conclusions of experimental investigations

From the small scale tests it can be derived:

1. There is a clear effect of the heat input on the maximum temperatures, cooling times t5/3 and the resulting deformations. Both maximum temperatures, cooling times t5/3 grow with increasing heat input per unit length. In principal, the deformations increase also with increasing heat input per unit length. But there exists a level for which the deformations reach a maximum. By further increase of the heat input per unit length, the deformations start to decrease. The reason for this behaviour lies within the reduced thermal gradient and the phase transformation in the region with the highest temperature. For plate thicknesses t = 50 mm and for high-strength steels (S690 and S890) this critical heat input per unit length could not be identified due to the restrictions in maximum temperatures.

2. Thickness of the plate has two different effects:

a. From the thermal side higher thicknesses need larger heat input than smaller ones. The heat conduction tends to be more three dimensionally the thicker the plate thickness is leading to smaller cooling times t5/3.

b. Due to the higher stiffness on the mechanical side the critical value of the heat input when the deformations starts to decrease with increasing heat input is located at higher heat input values for higher plate thicknesses.

3. In the small scale tests the steel grade has a large effect. With increasing strength the resulting deformations decreased in the free end tests. In the fixed end tests with increasing strength also the remaining internal forces increased.

4. The residual stresses arising from the flame straightening process resemble those of a welding process that means tensile stresses within the heated region and compressive stresses in the surrounding colder parts. Due to the more distributed heat flux distribution of the gas nozzles the region subjected to tensile stresses is wider. The maximum tensile stresses reach nearly the yield strength of the base material for each material investigated. For specimen with higher heat input and therefore higher maximum temperatures directly beneath the torch the tensile stresses are reduced and tend to compression. This may have a reason in the phase transformations which take place during the heating phase.

5. In view of the material impact investigations results of the small scale tests, it can be concluded that no significant changes in the mechanical behaviour have been observed for most of the analysed situations. Only the S235J0 for the 20 mm thickness plate exhibits a clear decrease in impact toughness resistance as a consequence of some remarkable microstructural changes. For this case and until new investigation results are available it is strongly recommended to limit the heat input during flame straightening processes particularly with low thickness components. In high strength steels (S690QL, S890QL) the bainitic tempered microstructure is partially lost in the heated area, transforming into a ferritic and perlitic formation with fine grain for both investigated thicknesses.

6. The conclusion of the specimen completely heated in the furnace is that for the penetrative heat straightening operations, the maximum temperature recommended in CEN/TR 10347 [3] should be followed.

7. Testing procedures have been developed to quantify the heat input of different nozzles and the corresponding heat flux distribution. The procedures are given in the guideline Appendix 2.

From the large scale tests the following points can be concluded:

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1. For the bending of beams either continuous (e.g. heat band on the flange) or discrete heating patterns (vee heats) are applicable. Continuous heating patterns lead to a continuous curvature while discrete heating patterns lead to a polygonal curvature.

2. The location of the heat pattern has an influence on the resulting deformation. The nearer the heat pattern is located to the middle of the beam the larger is the deformation that results from this pattern.

3. The localized heating patterns (like vee heat) proved to be effective when located near additional loads due to the additional compressive stresses at the location of the heating pattern. For members with high moment of inertia (IPE, HEA) the heating results tend to zero without additional restraint and higher deformations are only achievable by additional loads or restraints considering the beam length. Therefore relatively large displacements were obtained for e.g. the T member about the y-axis without an additional load.

4. As for the small scale tests also in the large scale tests an effect of the steel grade could be observed meaning that for most of the tests for higher steel grades the deformations are smaller compared to lower steel grades when using the same process.

5. The comparison of rolled and welded profiles has shown two tendencies:

- For IPE profiles about the strong axis the results for welded and rolled sections were approximately the same.

- For IPE profiles about the weak axis the results of welded were about 70% larger than those of the rolled sections.

6. Instabilities might be possible when using additional weights in a load controlled way. This may cause an additional displacement or, in the worst case, lateral torsional buckling. Therefore additional weights should be secured (e.g. by a crane) or path controlled restraints should be chosen.

7. For plate structures the influence of the material strength was obvious. For lower steel grades the deformations from both welding process and flame straightening process are larger than for higher steel grades. The line heat process proved to be successful without the use of any additional restraints for both plates without and with longitudinal stiffeners.

8. The heat spot patterns showed small effect on the out-of-plane deformation of stiffened plates. Without an additional restraint a deformation could not be induced. By the use of a mechanical jack a deformation was generated where the heat spots near the mechanical jack had the largest effect.

9. Distortions from fabrication process like rolling and welding have successfully been straightened. For out-of-square defects of profiles line heats on the web led to the desired straightening effect. Also line heats have been used on welded T-elements to straighten the deformations originated from the welding process.

10. By heating processes compressive stresses can successfully be induced at hot spot locations for crack initiation (e.g. LME).

11. The investigation of material properties on large scale specimens shows that for superficial heats at temperature ranges below 650-700ºC, no significant changes are expected in the mechanical behaviour of the materials. For penetrative heats with temperatures above 700ºC some minor changes in mechanical response were observed, except for S235 steel grades in which, as seen in small scale tests, some significant variations in toughness and ductility parameters can be expected. For all these reasons, it is recommended not to perform the flame straightening processes at temperatures above 700º C for penetrative heating patterns, particularly for the case of ordinary ferritic-perlitic steels. For short superficial heating, higher temperature could be applied according to CEN/TR 10347 [3].

12. The heating in the workshop was observed to be a very probabilistic process. The parameters

- adjustment of the torch

- distance of torch and element

- velocity of the torch (e.g. line heat)

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- heating times (e.g. vee heat, heat spot)

- oscillation parameters

showed due to the manual processing a relative large scatter. For example, the heating time of the flange of a vee heat has an average value of 82s with a standard deviation of 7,6s. This has to be considered in the view of a prediction of the flame straightening result, as the heat input is one of the major parameters.

5.2 Conclusions of numerical investigations

The structural elements have been simulated by FEM-techniques to analyse the time dependant thermal, metallurgical and mechanical processes in a way that is not possible only by experiments to get an inside view of the mechanisms during the flame straightening trials. The numeric models were calibrated by the small scale tests and the full scale trials. Three different FE-codes (ANSYS, ABAQUS and SYSWELD) were used to identify adequate simulation techniques for the flame straightening process.

By numerical simulations with SYSWELD plates have been simulated to extend the knowledge gained by the small scale tests. All tests with short holding time have been recalculated and the results showed good agreement with the experiments. The analysis of the line heat process showed that the formation of the plastic strains is a three-dimensional mechanism which can hardly be handled by analytical means. Therefore parametric studies have been performed to analyse the effect of the heat input, the plate thickness, the steel grade and the size and type of nozzle on the resulting temperature cycles and deformations.

The linear relationship between the heat input per unit length and the maximum surface temperature and the resulting deformation in terms of the rotation angle was confirmed also for other combinations of plate thickness and nozzle. Due to the lower surface temperatures with multi-flame rosebud nozzles a higher heat input per unit length can be applied than with single orifice nozzles and, therefore, the resulting rotation angles at the line heat can be increased. The limit value of the heat input per unit length leading to decreasing deformations with growing heat input, which was observed in the small scale tests, could be recalculated by numerical means also for other combinations of nozzle, plate thickness and steel grade.

By division of the heat input per unit length by the plate thickness and multiplication of the rotation angle the effect of the plate thickness can be eliminated and, therefore, the relation between heat input and deformation can be given by one line for one steel grade and type of heat flux distribution. By the results of the small scale tests this line could be normalized to one line for all steel grades. With this linear relationship it is possible to estimate the deformation of a line heat process on a plate structure.

Coupled thermo-mechanical transient FE analyses were performed on T-elements for two different heating patterns, for 10 and 20 mm thick plates and for two different material grades. With the simulations the deformations of the T-elements investigated in the frame of the experimental investigations could be recalculated and the results showed good agreement with the deformations of the test specimen.

Numerical simulations on beam structures with ABAQUS were performed in order to evaluate the influence of initial residual stress, the magnitude of external load, the size of heating pattern and the heating sequence on the final displacement. In general, the introduction of the residual stress in the numerical model results in lower final displacement. The consideration of residual stresses reduces the influence of the heating sequence (web-flange vs. flange-web). For vee heats with a large width and by application of high external loads the formation of a local plastic hinge at the location of the vee heat is possible and should be considered, because out-of-plane deformations of the flange and the web are possible especially for slender profiles. Up to the formation of the plastic hinge the deformation increases linearly with the applied external forces.

To derive prediction means for beam structures parametric studies have been performed with a component method. The piece of the beam surrounding the heating pattern is “virtually” cut out of the whole beam and the single vee heat is simulated using the flame model and process parameters of a realistic flame bending process. From the resulting deformation of this component the rotation of at the heating pattern is calculated. The deformation of the whole beam is calculated by superposition of the

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single deformations. These simulations proved that the relationship between the bending moment arising from external loads or restraints and the remaining rotation is linear, and that with increasing bending moment the deformation grows. This model has been calibrated by large scale tests and the results obtained were in good agreement with them. Moreover, the fact that an entire beam can be simulated using just two or three single simulations leads to a significant reduction in the amount of time required to simulate it, constituting one of the main advantages of this method.

To facilitate the calculation of the flame straightening process of complex structures a small guideline has been prepared concerning the numerical simulation. The guideline is attached to the report in Appendix 3.

5.3 Conclusions of analytical investigations

Based upon the experimental investigations and the numerical simulations the mechanisms of the heating patterns have been further investigated in the view of the prediction of the deformations of bar shaped structures. Two analytical models have been derived. With model 1 critical temperatures Tcrit,1 and Tcrit,2 can be calculated with which the beginning plastification in the colder part of the section can be set limits. The resulting deformations can be estimated by assumption that the complete thermal strains result in a plastic shortening of the vee heat and therefore in a rotation angle. By model 2 the plastic strains are derived using a fibre model of the section at different locations of the heating patterns. Assuming the hypothesis of Bernoulli and linear elastic – ideal plastic material behaviour the strain distribution is calculated in this way that the resulting stress distribution fulfils the equilibrium with the internal forces at the location of the heating pattern. Compared to the experimental results the analytical models give a good prediction when detailed information on geometry and sequence of the heating pattern and the temperature distributions are given.

The application of the derived analytical models is simple to use and quick possibility to predict the deformations from heating patterns on beam structures providing a relative well estimation for the critical temperatures and deformations. The application of the methods is presented in the guideline where also the procedures and means how to achieve the (critical) temperatures is shown.

5.4 General conclusion

Within the project the flame straightening process has been analysed with various methods. The existing knowledge given in the literature, research reports and instruction material has been reviewed. By experimental, numerical and analytical examinations this information has been enlarged and in view of a clarification of the mechanisms used to derive prediction means for the deformations from line heats and heating patterns on beams.

A flame straightening process is supposed to be successful, when two conditions are fulfilled.

1. No detrimental effect has taken place within the material resp. the mechanical properties of the material have not changed due to the heating process. This is supposed to be the case when a maximum allowable temperature is not exceeded. For this purpose a recommendation for the maximum temperatures has been elaborated for the different steel grades and heating types (superficial and full section heating) as a result of the experimental investigations. The recommended values are given in chapter 6.3 of the guideline (Appendix 2).

2. The aspired bending has been applied or the distortion from prior fabrication processes has been straightened. This aspect is addressed by the developed prediction means given in chapter 7.

The crucial point of a successful estimation of the heating process is to limit the various input parameters which often depend on the manual adjustment and handling of the torch. To achieve a good prediction of the results of a flame straightening process the control of temperature, heat input and heating time is of major concern as these are the input parameters for the estimation of the deformation. Procedures for the quantification of the heat input by a torch, the control of the temperature for superficial and penetrative heating patterns have been developed. The enlarged knowledge has been summarized and together with the prediction methods described in a guideline.

Even if the quantification of the heat input and the control of the temperature as input parameters lead to a small additional effort, by the derived prediction methods the flame straightening process becomes more effective and less critical, as no time consuming “try-and-error”-procedures have to be performed

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and the risk of a detrimental effect on the material has been reduced either by temperature control or prediction. Thus, a significant improvement of the procedures, the efficiency and the feasibility of flame straightening was achieved.

Nevertheless not all heating patterns and structural configurations could be investigated in the project. But the methods and procedures developed and applied within the project can be extended to other heating patterns and more complex structures.

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6 Guideline

By all partners a guideline has been elaborated for the presentation and clarification of the flame straightening process. The guideline is addressed to people in the workshop but also to engineers working in the process engineering.

The information gained through the literature review and the work within the work packages of the project has been summarized in a guideline giving an overview on the flame straightening process on different kinds of heating processes.

The guideline is divided into the following chapters:

Explanation of mechanism

Necessary equipment

Heating patterns, sequences and restraints

Adjustment of energy

Quantification and control of heat input

Avoiding an impact on the material

Prediction mean

Quantification of displacement

Certainty of success

Examples

Procedure

It also includes examples for the estimation of the flame straightening result by the derived prediction means for plate and beam structures. The guideline is attached to the report in Appendix 2.

Furthermore a small guideline has been prepared concerning the numerical simulation of the flame straightening process. The guideline is attached to the report in Appendix 3.

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7 Exploitation and impact of research results

The knowledge on the flame straightening process that was wide spread has been gathered and analysed within the project. By extensive experimental investigations and numerical calculations the mechanisms have been clarified for heating patterns of major importance and analytical prediction means have been derived for the easy and simple estimation of the these processes. The enlarged knowledge has been summarized and together with the prediction methods described in a guideline.

Even when the quantification of the heat input and the control of the temperature as input parameters lead to a small additional effort, by the derived prediction methods the flame straightening process becomes more effective and less critical, as no time consuming “try-and-error”-procedures have to be performed and the risk of a detrimental effect on the material has been reduced either by temperature control or prediction. Thus, a significant improvement of the procedures, the efficiency and the feasibility of flame straightening was achieved. The improved flame straightening-knowledge and improved –procedures results into a more precise geometry, less detrimental effects on the material, less energy input, quicker and healthier production and therefore leads to a better economic position avoiding expensive repair work.

To disseminate the knowledge and experience that has been gained within the project the consortium is preparing various publications in papers on national and international level (e.g. Steel Construction) and on national and international conferences.

Furthermore the project partners are in contact with steel constructors and European and national steel construction associations (ECCS, DSTV) to disseminate the results (final report and guideline) to steel constructors.

All project partners continue to work on the research topic and plan to submit a proposal for the further development and dissemination of the project results.

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8 List of references

[1] Rykalin, N.N.: “Berechnung der Wärmevorgänge beim Schweißen“, Berlin, VEB Verlag Technik, 1957

[2] ECCS, Manual on Stability of Steel Structures. Second ed. 1976: European Convention for Constructional Steel Structures.

[3] Guidance for forming structural steels in processing: DD CEN/TR 10347:2006

[4] R.Richard Avent, David J. Mukai and Paul F. Robinson, Effect of Heat Straightening on Material Properties of Steel, Journal of Materials in Civil Engineering, Vol 12, No 3, August 2000, pp. 188-195

[5] R.Richard Avent and David J. Mukai, Heat-Straightening Repairs of Damaged Steel Bridges, A Technical Guide and Manual of Practice,Report No FHWA-IF-99-004, Federal Highway Administration, October 1.1998

[6] R.Richard Avent, Heat-Straightening of Steel: From Art to Science, American Institute of Steel Construction

[7] R. Richard Avent and David J. Mukai, Engineered Heat-Straightening

[8] R.Richard Avent, David J. Mukai, Paul F. Robinson and Randy J. Boudreaux, Heat-Straightening Damaged Steel Plate Elements, Journal of Structural Engineering, Vol. 126, No 7, July 2000, pp. 747-754

[9] R.Richard Avent, Engineered Heat-Straightening, American Institute of Steel Construction

[10] Iowa Department of Transportation, Iowa Field Study Documented: Successful Heat-Straightening Repair of a Steel Bridge by In-House Personnel, http://www.usroads.com/journals/p/rmj/9710/rm971001.htm

[11] R. Richard Avent, Krishna K. Verma, Principles and Practice of Heat-Straightening Repair, American Institute of Steel Construction

[12] R. Richard Avent, David J. Mukai, and Ernest Heymsfield, Repair of Localized Damage in Steel by Heat Straightening, Journal of Structural Engineering, Vol. 127, No 10, October 2001, pp. 1121-1128

[13] R. Richard Avent, David J. Mukai and Paul F. Robinson, Residual Stress in Heat-Straightened Steel Members, Journal of Materials in Civil Engineering, Vol. 13, No 1, January/February 2001,pp. 18-25

[14] R. Richard Avent, David J. Mukai and Paul F. Robinson, Heat Straightening Rolled Shapes, Journal of Structural Engineering, Vol. 126, No. 7, July 2000, pp. 755-763

[15] Abdul Waheed, Ed Kowal and Tom Loo, Repair of Bridge Structural Steel Elements Manual, Bridge Engineering Section, Technical Standards Branch, Alberta Transportation, June 2004

[16] R.Richard Avent, Designing Heat Straightening Repairs, Proc., Nat. Steel Constr. Conf., American Institute for Steel Construction, Chicago, 1992

[17] R.L.Rothman, Flame Straightening Quenched-and-tempered steels in ship construction, Final Technical Report, Ship Structure Committee, 1974

[18] A.H.Varma, K. Kowalkowski, Effects of Multiple Damage Heat Straightening Repairs on the Structural Properties of Bridge Steels Final Report, Michigan State University, November 2004

[19] Pfeiffer,Richard:Handbuch der Flammrichttechnik: Metallurgie, Verfahren, Geräte und Anwendungsbeispiele, Fachbuchreihe Schweißtechnik Bd. 124, Deutscher Verlag f. Schweißtechnik, DVS-Verlag, Düsseldorf 1996

[20] Pfeiffer, Richard: Richten und Umformen mit der Flamme: Leitfaden für Ausbildung u. Praxis.2., überarbeitete u. erw. Auflage, Dt. Verl. für Schweißtechnik, Düsseldorf 1989

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[21] Dennin, G.: Einfluss des Flammrichtens auf das Verformungsverhalten und die Korrosionsbeständigkeit von Grobblechen aus unlegierten und legierten Stählen, Schw. Schn. 28 (1976), H. 11, S. 421/25

[22] Schärfer: Richten verzogener Bauteile mit dem Lichtbogen. Praktiker 24 (1972), H. 3, S. 252/53

[23] Anders, W.: Bewegungsablauf eines rechteckigen Stabes unter Einfluss eines wandernden Schutzgaslichtbogens in Abhängigkeit von der Stabhöhe, Schweißtechnik (Berlin) 16 (1966), H. 9, S. 389/94

[24] Pfeiffer, R.: Das Richten mit der Flamme. Sonderheft Schw. Schn. 4 (1952), S. 39/50

[25] Dennin, G.: Flammrichten von 5 bis 50 mm dicken Blechen aus verschiedenen Stählen, DVS-Berichte Bd. 57, S. 60/67, Deutscher Verlag f. Schweißtechnik DVS-Verlag, Düsseldorf 1979

[26] Merkblatt DVS 1602: Flammrichten der Beblechung von Schienenfahrzeugen, Teil 1: Stahl, Teil 2: Chrom-Nickel-Stahl, Teil 3: Aluminium, Deutscher Verlag f. Schweißtechnik DVS- Verlag, Düsseldorf

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[36] Thiele, W.-R.: Gefüge von hochfesten Feinkornbaustählen nach Flammrichtvorgängen. Schw. Schn. 36 (1984), H. 12, S. 579/83

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[38] Nieß, M., u. V. Schuler: Auswirkungen des Flammrichtens auf die mechanischen Gütewerte von hochfesten Feinkornbaustählen. DVS-Berichte Bd. 112, S. 169/73. Deutscher Verlag f. Schweißtechnik DVS-Verlag, Düsseldorf 1988

[39] Hanus, F. E.: Flammrichten thermomechanisch gewalzter Baustähle. Schw. Schn. 46 (1994), H. 4, S. 163/66

[40] Ehrenberg, H.: Aus der Praxis der Flammrichtens. Schweißtechnik(1986), H.9, S.190/93

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[44] [76] Engineered heat straightening comes of age, R.Richard Avent, Modern steel construction February 1995, p32-39

[45] Ph. Roguin C, Prince- F. Decaestecker : Description of a modus operanti for the execution of flame straightening with LINE pattern. (Institut de Soudure) From the report: Description d’un mode opératoire d’exécution des chaudes de retrait en ligne, pour la qualification des opérateurs ou la vérification des propriétés d’emploi du métal de base. Institut de soudure, rapport n° 30049

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[52] Hubo, Ralf; Kügler, Dietmar; Petersen, Jens; Wegmann, Horst - Flammrichten thermomechanisch gewalzter Schiffbaustähle, Schweißen und Schneiden 46(1994), H.6, p. 277-281

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[66] Jyri Outinen, Jyrki Kesti & Pentti Mäikeläinen: Fire Design Model for Structural Steel S355 Based Upon Transient State Tensile Test Results. J. Construct. Steel Res. Vol. 42, No. 3, pp. 161-169, 1997.

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9 List of figures

Figure 1-1: Flow chart of the flame straightening process......................................................... 15 Figure 1-2: Interrelationship of thermodynamics, mechanics and metallurgy........................... 15 Figure 1-3: Mechanism of line heat process on a plate (schematic) .......................................... 16 Figure 1-4: Work packages of the project and structure of the project ...................................... 16 Figure 2-1: Nozzles FB-A 8 to FB-A 10 (left side), detail of nozzle FB-A 9 (right side) ......... 19 Figure 2-2: effective heat input vs. distance of nozzle from plate surface................................. 20 Figure 2-3: Distribution of heat flux for selected nozzles .......................................................... 21 Figure 2-4: Test setup (left: view of thermo vision camera, right: position of thermo vision camera) ....................................................................................................................................... 21 Figure 2-5: left: Free end setup, right: location of thermocouples for both test setups ............. 22 Figure 2-6: Fixed end setup ........................................................................................................ 22 Figure 2-7: Maximum temperature vs. heating time .................................................................. 23 Figure 2-8: maximum temperature vs. velocity of the torch (left: t = 20 mm, right: t = 50 mm) .................................................................................................................................................... 23 Figure 2-9: cooling time t5/3 vs. velocity of the torch (left: t = 20 mm, right: t = 50 mm) ....... 24 Figure 2-10: out-of-plane deformation vs. heat input per unit length (left: t = 20 mm, right: t = 50 mm) ....................................................................................................................................... 24 Figure 2-11: temperature cycle of test S355_t20_free_sh_T3 (left: heating, right: cooling) .... 24 Figure 2-12: out-of-plane deformation cycle of test S355_t20_free_sh_T3 (left: heating, right: cooling)....................................................................................................................................... 25 Figure 2-13: out-of-plane deformation vs. yield strength and heat input................................... 25 Figure 2-14: temperature cycle of test S235_t50_free_sh_T3 (left: heating, right: cooling) .... 26 Figure 2-15: out-of-plane deformation cycle of test S235_t50_free_sh_T3 (left: heating, right: cooling)....................................................................................................................................... 26 Figure 2-16: out-of-plane deformation vs. yield strength and heat input................................... 26 Figure 2-17: temperature cycle of test S355_t20_free_lh_T3 (left: heating, right: cooling) ..... 27 Figure 2-18: out-of-plane deformation cycle of test S355_t20_free_lh_T3 (left: heating, right: cooling)....................................................................................................................................... 27 Figure 2-19: out-of-plane deformation vs. yield strength and maximum temperature .............. 28 Figure 2-20: development of normal force and bending moment in specimen S235_t20_fixeD_sh_T1 (left: heating, right: cooling) ............................................................... 28 Figure 2-21: normal force and bending moment vs. yield strength and heat input per unit length (left: normal force, right: bending moment)............................................................................... 29 Figure 2-22: comparison of the thermal cycle: a) furnace simulation and b) penetrative heat straightening trial........................................................................................................................ 29 Figure 2-23: Tensile properties (Re, Rm) after heat simulation in the furnace ........................... 30 Figure 2-24: Yield stress (left) and tensile strength (right) in the specimens obtained from the 20 mm thick plate ....................................................................................................................... 31 Figure 2-25: Yield stress (left) and tensile strength (right) in the specimens obtained from the 50 mm thick plate ....................................................................................................................... 31 Figure 2-26: Charpy curves (absorbed energy vs. temperature) for S235J0 steel. Left, 20 mm plate; right, 50 mm plate. ........................................................................................................... 32 Figure 2-27: Charpy curves (absorbed energy vs. temperature) for S355J2 steel. Left, 20 mm plate; right, 50 mm plate. ........................................................................................................... 32 Figure 2.28: Charpy curves (absorbed energy vs. temperature) for S460ML. Left, 20 mm plate; right, 50 mm plate. ........................................................................................................... 32 Figure 2-29: Charpy curves (absorbed energy vs. temperature) for S690QL steel grade. Left, 20 mm plate; right, 50 mm plate. .................................................................................................... 33 Figure 2-30: Charpy curves (absorbed energy vs. temperature) for S890QL steel. Left, 20 mm plate; right, 50 mm plate. ........................................................................................................... 33 

105

Figure 2-31: Micro structure for S235J0 taken from 20 mm thickness plate. Left, as-received state; right, after flame straightening. ......................................................................................... 34 Figure 2-32: Summary of the results of micro-hardness obtained for 20 mm (left) and 50 mm (right) .......................................................................................................................................... 34 Figure 2-33: Microstructure for ferritic-perlitic steels (S235, S355 and S460) as-received material taken from 20 mm thickness. ....................................................................................... 35 Figure 2-34: Microstructure for ferritic-perlitic steels (S235, S355 and S460) heated material taken from 20 mm thickness plate ............................................................................................. 35 Figure 2-35: Microstructure for ferritic-perlitic steels (S235, S355 and S460) as-received material taken from 50 mm thickness ........................................................................................ 35 Figure 2-36: Microstructure for ferritic-perlitic steels (S235, S355 and S460) heated material taken from 50 mm thickness ...................................................................................................... 35 Figure 2-37: Microstructure for high strength steels (S690 and S890) as-received material taken from 20 mm thickness plate ............................................................................................. 36 Figure 2-38: Microstructure for high strength steels (S690 and S890) heated material taken from 20 mm thickness plate ....................................................................................................... 36 Figure 2-39: Microstructure for high strength steels (S690 and S890) as received material taken from 50 mm thickness plate ....................................................................................................... 36 Figure 2-40: Microstructure for high strength steels (S690 and S890) heated material taken from 20 mm thickness plate ....................................................................................................... 36 Figure 2-41: Distribution of strain gauges and cutting lines ...................................................... 37 Figure 2-42: Water jet cutting of 50 mm thick plate (left), plate after cutting........................... 37 Figure 2-43: Measured residual stresses for 20 mm thick plate ................................................. 38 Figure 2-44: Measured residual stresses for 50 mm thick plate ................................................. 38 Figure 2-45: residual longitudinal stress for test S235_t20_free_sh_T3 ................................... 39 Figure 2-46: residual longitudinal stress for test S355_t20_free_sh_T3 ................................... 39 Figure 2-47: heat straightening operations to act on the flange parallelism .............................. 40 Figure 2-48 : Example of location and sequence of heating pattern along the strong axis ........ 42 Figure 2-49: Example of location and sequence of heating pattern along the strong axis ......... 42 Figure 2-50: Band pattern investigated to induce a bending along weak axis ........................... 43 Figure 2-51: Location and sequence to induce a bending along the strong axis on L profiles .. 43 Figure 2-52: Location and sequence to induce a bending along the strong axis on T profiles .. 43 Figure 2-53 Location and sequence to induce a bending along the weak axis on T profiles ..... 43 Figure 2-54: record of vertical displacements for test 4.1-13 .................................................... 44 Figure 2-55: location of deformation transducers and heating pattern of test 4.2-1 and 4.2-4 .. 45 Figure 2-56: resulting out-of-plane deformation for test 4.2-1 .................................................. 45 Figure 2-57: Location of Deformation Transducers (DT), Field 1 ............................................ 45 Figure 2-58: Field 1of test 4.2-2 with lifting jack ...................................................................... 45 Figure 2-59 Heating patterns for the three fields of test 4.2-2 ................................................... 46 Figure 2-60: Heating patterns for the three fields of test 4.2-5 .................................................. 46 Figure 2-61: Deformation record of field 1of test 4.2-2 ............................................................ 46 Figure 2-62: heating pattern of test 4.2-3 ................................................................................... 47 Figure 2-63: Out of plane deformation of test 4.2-3 (S235) ...................................................... 48 Figure 2-64: Heat pattern to straighten an out-of-square on one flange (data sheet 4.3-11) ..... 49 Figure 2-65 : Heating pattern to straighten a similar out-of-square on the two flanges on a section......................................................................................................................................... 49 Figure 2-66: description of the transversal bands pattern (second heat performed for trial 4.3-13)............................................................................................................................................... 50 Figure 2-67: Description of the second heating pattern (band pattern) applied for the trial 4.3-14 ................................................................................................................................................ 51 Figure 2-68: Distortions induced by the welding for T and plates assemblies .......................... 52 Figure 2-69: Heating pattern on free ends .................................................................................. 54 

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Figure 2-70: Heating pattern on detail half end cover plate ....................................................... 54 Figure 2-71: Heating pattern on detail cope cut ......................................................................... 54 Figure 2-72: Transversal residual stresses, free end detail, S235, IPE 550 ............................... 55 Figure 2-73: Transversal residual stresses, free end detail, S460, IPE 550 ............................... 55 Figure 2-74: Transversal residual stresses, cope cut, IPE 550 ................................................... 55 Figure 2-75: Transversal residual stresses, half end cover plate, S460, IPE 550....................... 55 Figure 2-76: Location of residual stress measurement for test 4.1-15 ....................................... 58 Figure 2-77: Residual stresses in the flanges of test 4.1-15 (IPE 450, S355J2, vee heat n° 11 about the strong axis) ................................................................................................................. 59 Figure 2-78: Residual stresses in the web of test 4.1-15 (IPE 450, S355J2, vee heat n° 11 about the strong axis) ........................................................................................................................... 59 Figure 2-79: Location of residual stress measurement for test 4.1-16 ....................................... 59 Figure 2-80: Residual stresses in the flange of test 4.1-16 (IPE 450, S355J2, vee heat n° 4 about the weak axis) ................................................................................................................... 59 Figure 3-1: Distribution of the heat flow density according to equation (3.5)........................... 61 Figure 3-2: Effective heat input and degree of efficiency vs. acetylene consumption .............. 62 Figure 3-3: Maximum heat flow density and effective diameter of heat source vs. acetylene consumption ............................................................................................................................... 62 Figure 3-4: Temperature over depth of the plate for different finite element types at t = 39 s .. 63 Figure 3-5: Stresses over depth of the plate for different finite element types .......................... 64 Figure 3-6: Temperatures during the heating process (t = 120 s) .............................................. 65 Figure 3-7: Temperatures during the heating process (t = 120 s) .............................................. 65 Figure 3-8: Stresses in transversal direction during the heating process (t = 120 s) .................. 65 Figure 3-9: Stresses in longitudinal direction during the heating process (t = 120 s) ................ 65 Figure 3-10: stresses in transversal direction during the heating process (t = 120 s) ................ 65 Figure 3-11: Stresses and plastic strains in transversal direction along the centre line during the heating process (t = 120 s) ......................................................................................................... 66 Figure 3-12: Development of plastic strains in during the heating process ............................... 66 Figure 3-13: maximum temperature vs. heat input per unit length ............................................ 66 Figure 3-14: rotation angle vs. maximum temperature .............................................................. 66 Figure 3-15: rotation angle vs. heat input per unit length .......................................................... 67 Figure 3-16: thickness x rotation angle vs. heat input per unit length and thickness ................. 67 Figure 3-17: thickness x rotation angle vs. heat input per unit length and thickness for different steel grades ................................................................................................................................. 67 Figure 3-18: thickness x rotation angle / m vs. heat input per unit length and thickness .......... 67 Figure 3-19: Definition of heating areas .................................................................................... 68 Figure 3-20: Surface temperature at 45 s (element T5) ............................................................. 68 Figure 3-21: The effect of heating time, power of the burner on the displacement and surface temperature for element T5 ........................................................................................................ 69 Figure 3-22: Basic input data and experimental results for element T5 .................................... 69 Figure 3-23: Comparison of experimental and FEA results for element T5 .............................. 70 Figure 3-24: Comparison of experimental and FEA results for element T2 .............................. 70 Figure 3-25: The basic input data and the final displacement for element T6 for the experiment .................................................................................................................................................... 71 Figure 3-26: The effect of heating time, power of the burner on the displacement and surface temperature for element T6 ........................................................................................................ 71 Figure 3-27: The positions of V heats on the beam ................................................................... 71 Figure 3-28: The numerical model of burners ........................................................................... 71 Figure 3-29: The temperature dependent stress-strain curves as input data for numerical simulations ................................................................................................................................. 72 Figure 3-30: The initial residual stress in the flange and web of IPE450 according to the literature [2] (ECCS) and as an input data in the numerical analyses (Abaqus) ........................ 72 

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Figure 3-31: The temperature in point A and B only for the heating phase for analyses S1, S1RS, S5 and S5RS .................................................................................................................... 73 Figure 3-32: Vertical displacement (the influence of the initial residual stress and heating sequence) .................................................................................................................................... 73 Figure 3-33: The deflection of the beam in heating and cooling phase and the final deflection (the influence of the initial residual stress, width of V heat and external load) ......................... 73 Figure 3-34: Final vertical displacements in dependence to the external force for different widths of V heat ......................................................................................................................... 74 Figure 3-35: Equivalent plastic strain after unloading (Deformation Scale Factor = 10 ) in the mid-section point ........................................................................................................................ 74 Figure 3-36: The development of the vertical displacement in coupled steady-state thermomechanical analysis ........................................................................................................ 76 Figure 3-37: Normal stress component in the cross-section at different temperatures of V heat (height of V heat form z = –225 to z = 75) ................................................................................ 76 Figure 3-38: Simplification of an entire beam into several short beams ................................... 77 Figure 3-39: Experimental verification for the thermal independency over a steel plate. ......... 77 Figure 3-40: Sketch of the estimation of the rotation angle in the short beam. ......................... 78 Figure 3-41: Final deflection sketch........................................................................................... 78 Figure 3-42: HEA300 Bending moment-Rotation graph ........................................................... 78 Figure 3-43: Heating patterns, discretised area .......................................................................... 79 Figure 3-44: HEA 300 S355 test setup....................................................................................... 79 Figure 3-45: HEA 300 S460 test setup....................................................................................... 79 Figure 3-46: IPE 450 S460ML test setup ................................................................................... 80 Figure 4-1: Deformation of the beam due to plastic strain in V heat ......................................... 82 Figure 4-2: Simplified stress distribution in the V heat at different temperature levels ............ 82 Figure 4-3: Ratio of Young’s Modulus vs. temperature ............................................................ 84 Figure 4-4: Ratio of yield strength vs. temperature ................................................................... 84 Figure 4-5: coordinate system and internal forces ..................................................................... 84 Figure 4-6: Thermal strains ........................................................................................................ 84 Figure 4-7: Stress distribution .................................................................................................... 85 Figure 4-8: Comparison of analytical and experimental results for tests 4.1-13 ....................... 87 Figure 4-9: Comparison of analytical and experimental results for tests 4.1-14 and 4.1-17 ..... 87 Figure 4-10: strain and stress distributions (IPE 450 profile, S235, vee heat: 300 mm/80 mm) 88 Figure 4-11: Distribution of in a vee heat ............................................................................... 88 Figure 4-12: Distribution of for a longitudinal line heat at the centre of the flange ............... 88 Figure 4-13: Temperature distribution on the surface of the flange .......................................... 89 Figure 4-14: Temperature distribution on the surface of the web .............................................. 89 Figure 4-15: Comparison of analytical and numerical rotations in relation to the temperature in the vee heat ................................................................................................................................. 90 Figure A2-5.1: Steel plate (1) flame and thermocouples position ........................................... 131 Figure A2-5-2: Steel plate (2) flame and thermocouples position ........................................... 131 Figure A2-5.3: Temperature results in plate 2 ......................................................................... 132 Figure A2-5.4: Curves for adjusting flame parameters ............................................................ 132 Figure A2-6.1: Delivery conditions ......................................................................................... 133 Figure A2-6.2: Influence of increasing tempering temperatures on the tensile properties (left) and on the Charpy V transition temperature (right) - S890QL, 60 mm ................................... 134 Figure A2-6.3: Microstructures of various delivery conditions ............................................... 135 Figure A2-7.1: maximum temperature vs. heat input per unit length ...................................... 137 Figure A2-7.2: thickness x rotation angle / m vs. heat input per unit length and thickness .... 138 Figure A2-7.3: Flame straightening process of a plate with T-stiffeners (left: performance, right: location of line heats (FS1 to FS 7) ................................................................................ 138 Figure A2-7.4: Deformation record for line heat FS3 .............................................................. 139 

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Figure A2-7.5: Deformation record for line heats FS5 and FS6 .............................................. 139 Figure A2-7.6: Records of the thermo vision camera ( = 0,85) of the line heats (left: FS3, middle: FS5, right: FS6) ........................................................................................................... 140 Figure A2-7.7: Reduction factor for the proportional stress at elevated temperature .............. 140 Figure A2-7.8: Geometrical factor kg for IPE, HEA, HEB, HEM profiles for full depth of V heat ........................................................................................................................................... 141 Figure A2-7.9: Rotation of the V heat at the critical temperature for IPE 400; material model S235 (mean values); full depth of V heat ................................................................................. 142 Figure A2-7.10: Rotation of the V heat at the critical temperature for IPE 400; material model EN 1993-1-2; full depth of V heat ........................................................................................... 142 Figure A2-7.11: Rotation of the V heat at the critical temperature for angle of V heat 30° and S235 (mean values) ; full depth of V heat ................................................................................ 142 Figure A2-7.12: HEA 300 temperature-rotation angle graph .................................................. 143 Figure A2-7.13: IPE 450 temperature-rotation angle graph .................................................... 144 Figure A2-7.14: HEM 340 temperature-rotation angle graph.................................................. 144 Figure A2-7.15: HEA 300 bending moment- rotation angle graph ......................................... 144 Figure A2-7.16: IPE 450 bending moment- rotation angle graph ............................................ 145 Figure A2-7.17: HEM 340 bending moment- rotation angle graph ......................................... 145 Figure A2-7.18: Rotation as a function of the steel grade ....................................................... 145 Figure A2-7.19: Test setup. The first step in red, the second in blue ...................................... 146 Figure A2-7.20: Flame straightening process of the profile. .................................................. 146 Figure A2-7.21: Test setup. Thermocouples and vees and band shape ................................... 146 Figure A2-7.22: Temperature and deflection over time ........................................................... 147 Figure A2-7.23: Test setup. ...................................................................................................... 147 Figure A2-7.24: Flame straightening process of the profile .................................................... 148 Figure A2-7.25: Thermocouples position and shape of the V ................................................. 148 Figure A2-7.26: Temperature and deflection over time ........................................................... 148 Figure A2-7.27: strain and stress distributions (IPE 450 profile, vee heat: 300 mm/80 mm) . 149 Figure A2-7.28: strain and stress distributions (IPE 450 profile, S235, vee heat: width 80 mm) .................................................................................................................................................. 149 Figure A3-6.1: Simplification of the entire beam in short beams ............................................ 152 Figure A3-6.2: Meshed Model of short beam .......................................................................... 152 Figure A3-6.3: Rotation angle in a short beam ........................................................................ 152 Figure A3-6.4: Final deflection sketch ..................................................................................... 153 

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10 List of tables

Table 2-1: Heat input for nozzles N°9 and N°10 ....................................................................... 20 Table 2-2: Chemical composition of plates (% mass)................................................................ 31 Table 2-3: Synthesis of the results of mechanical properties afterheat straightening ................ 41 Table 2-4: Summary of the flame parameters and restraint load used for the trials 4.1-1 to 4.1-43 ................................................................................................................................................ 42 Table 2-5: Deformations for test 4.2-2 (S235) ........................................................................... 46 Table 2-6: Deformations for test 4.2-5 (S460) ........................................................................... 47 Table 2-7: The test protocol and results of flame straightening of T elements (four measuring points for each T element) .......................................................................................................... 53 Table 2-8: Parameters of the temperature for ends of profiles ................................................... 54 Table 2-9: Chemical composition of investigated large scale specimen (wt-%) ....................... 56 Table 2-10: Large scale specimens selected for material impact investigation ......................... 56 Table 2-11: Results of Tensile and Charpy test on material from large scale specimens .......... 57 Table 3-1 : Thermal specifications for common burner nozzles ................................................ 62 Table 3-2 : heat input parameters for nozzles investigated within the project ........................... 63 Table 3-3 : Slope m for different steel grades (average values)................................................. 67 Table 3-4: Input data and final displacement in numerical simulations of IPE450 ................... 75 Table 3-5: Summary of the results. ............................................................................................ 79 Table 3-6: Summary of the results. ............................................................................................ 80 Table 3-7. Summary of the results. ............................................................................................ 80 Table A1-1: Documents prepared for meetings ....................................................................... 115 Table A1-2: Working documents and 6month reports ............................................................. 116 Table A2-2.1: Recommended torch tip for various thickness material. ................................... 119 Table A2-5.1 : heat input parameters for nozzles investigated within the project ................... 130 Table A2-7.1 : Slope m for different steel grades (average values) ......................................... 137 Table A2-7.2 : Prediction of the deformation of the considered line heats ............................. 138 Table A2-7.3: Summary results ............................................................................................... 147 Table A2-7.4: Summary of results ........................................................................................... 148 

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11 List of acronyms and abbreviations

Abbreviation Description unit

Young’s Modulus (Index RT = room temperature) MPa, N/mm²

N Normal force N, kN

My, Mz Bending moment kNm

Stress MPa, N/mm²

m,i Mechanical strain, i = location -

pl,i Plastic strain, i = component -

y Strain at yield strength -

th, T Thermal strains -

th, T Thermal expansion coefficient 1/K

Emissivity -

Q Energy J

q Heat input, Energy per time, effective value J/s

qu Total heat input J/s

qs Heat input per unit length, q/v J/mm

Heat input per unit length and plate thickness, q/v J/mm²

dV/dt Gas flow l/h, l/min

Maximum heat flux J/mm²s

r0,05 radius of the heat source mm

k Concentration factor of the Gaussian distribution 1/mm²

v Velocity of the torch mm/s

m Mass kg, t

T Temperature ° C

T Temperature difference K

Ambient temperature ° C

Tmax Maximum temperature ° C

Tcrit,i Critical temperature ° C

c Specific heat of a material J/kgK

Degree of efficiency -

m Slope rad mm³/J

fu Tensile strength MPa, N/mm²

fy Yield strength (Index RT = room temperature) MPa, N/mm²

fp Proportionality limit MPa, N/mm²

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Mvee moment due to external force at x = lv, kNm

Ac area of the cross section in the compression, mm²

cT the distance between centres of gravity of the cross-section in compression and in tension.

mm

curvature 1/mm

Rotation angle rad

t time s

t5/3 Time difference in a cooling cycle between 500° C and 300° C s

theating, th Heating time s

t Plate thickness mm

b Width of a profile mm

h Height of a profile mm

bvee Width of a vee heat mm

hvee Height of the vee heat mm

vee Angle of the vee heat °

LHV Lower heating value J/l, kJ/kg

AR As rolled -

N Normalised -

Q quenched and tempered -

M thermo mechanically rolled -

, f, w Deformations mm

Wpl Plastic section modulus mm³

kg geometrical factor -

kf load factor -

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Appendix 1: List of documents distributed in the frame of OPTISTRAIGHT

The following documents were prepared by the project members during the project:

Table A1-1: Documents prepared for meetings

Number Date Location Minutes Presentations

01 18-10-2007 Aachen Agenda Minutes

RWTH-1 RWTH-2 AGDH-1 APLR-1 APLR-2 UCAN-1 UCAN-2

02 07-02-2008 Esch-sur-Alzette Agenda Minutes

RWTH-1 RWTH-2 APLR-1 APLR-2

UL-1

03 25-09-2008 Dillingen Agenda Minutes

RWTH-1 RWTH-2 APLR-1 APLR-2

UL-1 UL-2

UCAN-1 UCAN-2

04 12/13-02-2009 Santander Agenda Minutes

RWTH-1 RWTH-2 RWTH-3 RWTH-4 APLR-1 APLR-2

UL-1 UCAN-1 UCAN-2

05 24/25-09-2009 Ljubljana Agenda Minutes

RWTH-1 RWTH-2 APLR-1 APLR-2

UL-1 UL-2 UL-3

UCAN-1

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Number Date Location Minutes Presentations

06 04/05-03-2010 Aachen Agenda Minutes

RWTH-1 RWTH-2 RWTH-3 APLR-1

UL-1 UL-2 UL-3

UCAN-1 UCAN-2

07 23/24-09-2010 Ljubljana Agenda Minutes

RWTH-1 RWTH-2 APLR-1

UL-1 UL-2 UL-3

UCAN-1 UCAN-2

Table A1-2: Working documents and 6month reports

Number Date Topic 01 07.2007 Project proposal RFS-PR-06067 02 07.2007 Project contract RFSR-CT-2007-00040 03 03.2008 Six Monthly Technical Report No 1 04 09.2008 Six Monthly Technical Report No 2 05 03.2009 Mid-term technical implementation report

06 05.2009 Presentation of Mid-term technical implementation report to the TGS8-Committee

07 09.2009 Six Monthly Technical Report No 3 08 03.2010 Six Monthly Technical Report No 4 09 10.2010 Draft Final report 10 10.2010 Word template for formatting 11 10.2010 Cost statement forms for final financial statements

Furthermore the documentation of the large scale tests and the tests results are available for download at the project web page:

http://www.stb.rwth-aachen.de/projekte/2007/OPTISTRAIGHT/OPTISTRAIGHT.html

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Appendix 2: Guideline

The following guideline shall give an overview on the flame straightening process including the results gained within the project OPTISTRAIGHT. The guideline is addressed to people in the workshop but also to engineers working in the process engineering.

1 Explanation of mechanism

Shape control by local heating is often used in processing of structural steels to correct deformations which occur during production or welding of the plates/beams. Experience has proven the value of this operation in many applications and for many different steel types.

The flame straightening of a plate / beam is based on a local heating of the area to be shortened in combination with the hindrance of thermal expansion by the cold vicinity.

This causes a local plastic compression in the heated zone. When the area subsequently cools down to ambient temperature, the thermal shrinking finally results in residual elastic tensile stresses that lead to a change of the shape. A plastic deformation that would compensate the initial compressive straining (at the hot spot) is not happening because the yield strength of the material has considerably increased at the lower temperature.

The efficiency of the thermal shaping process depends on

- form of the torch, distance from the surface, heating gas, gas flow rate - energy input - thermal field - heating rate - wall thickness and size of the cross section to be shaped - heating pattern (spot, line, triangle) - restraint and stiffness of the construction

To ensure a very rapid heating of a limited area of the plate / beam Acetylene shall be preferred due to the higher flame temperature (3200°C) compared to natural gas (2700°C)

During heating the structure will deform into the wrong direction. Only at the late stage of cooling the shape will be improved compared to the initial situation. The straightening effect can only be measured after the structure has completely cooled down and has attained an even temperature. The cooling can be accelerated by water when the temperature of the element has reached values below 315° C.

If during heating the structure is restraint by reinforcement or by applying external loads, so that the compressive straining is enhanced, the effect of shaping will increase.

At peak temperature the area to be straightened must undergo plastic deformation. The final tensile stresses are producing the effect on geometry. Thus the peak temperature must be sufficient to produce:

- Sufficiently high compressive stress at the hot spot. - The yield strength at elevated temperature is lowered to such an extent that the compressive

stress will lead to local plastic deformation.

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- The depth of plastically deformed material relative to the plate thickness must be sufficient to produce (macroscopic) bending.

For structural steel and beams both conditions are fulfilled for temperatures exceeding approximately 450°C. Any lower temperature will be ineffective for shaping unless mechanical devices are increasing the compressive stresses in the hot area.

The effect of the material strength has to be taken into account. Higher steel grades have higher yield strength even at higher temperatures which results in the effect that the deformations from the same flame straightening process are smaller for steels with higher strength.

2 Equipment, adjustment of energy and oxygen partition of the torch

2.1 Heat input

2.1.1 Gas

2.1.1.1 Acetylene

The Acetylene fuel is preferred by many users because it is a “hot” fuel. However, this fuel is also highly volatile. The transference of the heat to the material is better than the propane flame. The temperature in the flame is up to 3200ºC. As a guideline, it is recommended to use a mixture of 1 part of acetylene, 1.05-1.1 parts of oxygen and a pressure of 1.1 bar, oxygen 5.5-6 bar. These recommended values have been collected from experimental experience during the laboratory tests at the workshops.

2.1.1.2 Propane

The propane flame is safer to handle. Since it does not burn as hot, a larger tip or rosebud orifice may be required. The maximum temperature in the flame is around 2700 ºC. A mixture of 3:1 oxygen-propane is recommended; the pressure can be around 6 bars for the oxygen and 1.2 bars for propane. These recommended values have been collected from experimental experience during the laboratory tests.

2.1.2 Torch typology

The torch is tool for the heating process. Typically it is handled and adjusted manually. With most of the torch tips available on the market a large range of adjustment is possible. The adjustment depends on the individual experience of the workman which may lead to a large scatter in heat input from process to process.

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2.1.2.1 Single orifice torch tip

This type of torch consists of a brazing tip with a single opening for supplying fuel to the flame. The amount of heat applied is a function of the size of the tip. A larger size orifice provides a greater amount of heat than a smaller size orifice. Typically the efficiency decreases with increasing size of the torch.

Those tips, which are normally used for gas welding, generate a concentrated heat flux which leads to very high temperatures within a short time, especially with small plate thicknesses. Therefore an oscillation of the torch is recommended to distribute the heat flux over a larger area of the element.

2.1.2.2 Multiple orifice torch tip

This type of torch consists of a multiple orifice torch tip supplying fuel to the flame. The sizes of the orifices as well as the number of them determine the amount of heat applied by the torch. Larger orifices provide more heat, as do a higher number of orifices. The heat flux of those tips is less concentrated than for the single orifice tips.

The measure of the heating tips normally used goes from size nº3 to nº9 using oxyacetylene gas and nº 6 to nº10 using propane fuel. In Table A2-2.1 some recommendations for the selection of the type and size of the orifice as a function of the plate thickness can be found.

Steel Thickness (mm) Orifice Type Size 6 Single 3

10 Single 4 13 Single 5 16 Single 7 20 Single 8 25 Single 8 25 Rosebud 3 50 Single 8 50 Rosebud 4 75 Rosebud 5

100 Rosebud 5

Table A2-2.1: Recommended torch tip for various thickness material.

2.2 Measurement equipment

2.2.1 Temperature control

One of the most important and yet difficult-to-control parameters of heat straightening is the temperature of the heated metal. Factors affecting the temperature include size and type of the torch orifice, intensity of the flame, speed of torch movement, and thickness and configuration of the member. A summary of the main temperature control devices is presented below.

2.2.1.1 Thermocouples

A thermocouple is a versatile temperature sensor. The temperature is measured by contact between the thermocouple and the surface to be measured. They are inexpensive and interchangeable. The accuracy of these sensors is sufficient for the range of temperature of the heating treatments. However, because the thermocouple relies on full contact with a smooth surface, the readings vary with position and pressure, typically underestimating the actual temperature. They have very low inertia. Taking into account the experience collected in the developed large scale laboratory tests it is recommended to use the thermocouples type k (chrome-aluminium) for high temperatures. The thermocouple consists of bimetallic tip follow by a plastic protection. The plastic protection must be away from the direct action of the flame. After several uses the thermocouple must be replaced. Its working range is -200ºC to 1600ºC and its accuracy is 1ºC, 10ºC.

2.2.1.2 Contact thermometers

There are many types of these devices. The range of work for these devices goes from -200ºC to 1800ºC. These devices can be difficult to manipulate when the temperatures go too high. The accuracy of these devices for the temperature range of heating treatments usually goes from 20ºC to 40ºC.

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2.2.1.3 Optical and infrared pyrometers

Infrared and optical pyrometers allow temperatures to be measured in applications where conventional sensors cannot be employed or where non-contact measurements are required because of contamination or hazardous reasons or where the temperatures to be measured are too high for thermocouples or other contact sensors. These devices record the temperature and provide a digital readout. The critical considerations for any infrared pyrometer include field of view (target size and distance), type of surface being measured (emissivity considerations), spectral response (for atmospheric effects or transmission through surfaces), temperature range and mounting. They do not work appropriately on brilliant metallic surfaces and their working range is -40 ºC to 4000 ºC.

For correct adjustment the knowledge of the emissivity is necessary. In most cases an emissivity of = 0,80 – 0,85 is appropriate.

2.2.1.4 Thermographic cameras

These devices show an image of the entire area which is being heated. This feature allows more information than the other devices to be registered as the other techniques only provide information from one point. These devices are usually more expensive than the others and their working range is -20ºC to 900ºC and accuracy +-2%.

For correct adjustment the knowledge of the emissivity is necessary. In most cases an emissivity of = 0,80 – 0,85 is appropriate.

2.2.1.5 Temperature crayons

These crayons are manufactured with a material that melts and becomes liquid when their specified temperatures are exceeded. Also the crayons can be reapplied due to the simple nature of their use. There are lots of them and they are relatively inexpensive. The major uses are where a quick check of the temperature of an object is desired. The crayons do not allow the temperature to be measured directly - the user can only know when the material temperature rises above the specific temperature of the crayon. However, the flame tends to distort the results by blackening the crayon marks. To solve this problem, marks can be placed on the back side, although this might not allow the operator to see the results during the process. These crayons are available in increments of their specific temperature as small as 14°C. By using two crayons that bracket the desired heating temperature, an accurate control can be maintained. The crayons will burn if exposed directly to the flame of the torch. Therefore, the torch must be momentarily removed (one or two seconds) so that the crayons may be struck on the surface. Their working range is 120 to 600 ºC and their accuracy is +-5ºC.

2.2.1.6 Visual control

A very common method to control the temperature in steels is to check the surface colour. This procedure is highly dependent on the lighting conditions and the surface of the steel element.

Typically in normal daylight or interior lighting conditions, a 650ºC temperature will be indicated by a satiny silver colour near the torch tip. After cooling, the area should be grey in colour. A cherry red colour during heating or a black colour after cooling indicates that the heat was too high. However, this is the least accurate method of monitoring temperature and is not recommended.

2.2.2 Displacement control

2.2.2.1 Simple measuring tool

Simple measuring tools that are available in every workshop (e.g. measuring tape and rope/wire to provide a reference line) are accurate enough to control the displacements in the heat straightening process for most cases.

2.2.2.2 Inductive displacement transducer

This type of transducer needs to be in direct contact with the surface which displacement needs to be verified. This fact can be a problem in the heated areas because these devices are not usually designed for working under high temperature conditions. This requires taking additional precautions if the surface to be measured is heated up. These devices are highly precise.

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2.2.2.3 Laser displacement transducer

These devices base their way of working on the reflection of the laser light on the surface that is to be measured. To measure the displacement of a surface, these do not have to be in contact due to the nature of their use which is a great advantage when the surface to be measured is at a high temperature. In some cases, the laser light cannot reflect correctly but it generally works properly with steels. They are highly precise and easy to manipulate.

2.3 Additional loads or restraints

The deformation to be induced or the distortion to be straightened governs the choice of type, location and sequence of the heating pattern, the heat input and the heating times. From the process parameters results a transient temperature field in the region of the heating pattern. Of high importance for the resulting deformation is the generation of a thermal gradient over the thickness for superficial heating patterns like line heats or within the section between heated and non-heated part for penetrative heating patterns like vee heats. This gradient leads in combination with the material and structural configuration to an obstruction of the thermal strains arising from the temperature field. This obstruction results in the formation of negative plastic strains in the regions of high temperatures.

In cases of high stiffness of the structural element or if the heated part of the section is small compared to the remaining colder part the heating will not lead to the desired effect. By the application of adequate additional restraints or loads the formation of plastic strains during the heating phase can be enforced.

In contrast to the straightening with a flame during the mechanical straightening process external forces are applied to exceed the yield strength of the material at room temperature locally and, by this way, to impose a plastic strain state which leads after unloading to the desired straightening resp. deformation.

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In contrast to the additional loads or restraints for the flame straightening process, which shall not exceed 25% of plastic moment resistance of the section, the loads during mechanical straightening have to larger than the plastic moment resistance to induce a plastic deformation.

2.3.1 Mechanical jack

This type of jack is controlled by the user by activating mechanical actuators. These jacks are easy to use. It is necessary to have a rigid point of reaction to apply the force, and while the flame is applied and the structure is moving, the user needs to readjust the force constantly to have a constant applied force. Controlling the applied force with this type of jacks can be complicated if the deformations in the structure during the heating treatment are large.

2.3.2 Hydraulic jack

The only difference between the hydraulic and the mechanical jack lies in the way of applying the force. Like the mechanical jack, the hydraulic jack needs to be adjusted while the structure is moving to have a constant applied force.

2.3.3 Dead loads

This method consists in hanging down from the structure a dead load. With the dead load the applied force is always constant and in every moment of the heating treatment the operator know the exactly applied force however these loads are usually very heavies and their manipulation is difficult moreover obviously this method can be used only in the gravitational direction.

2.3.4 Location of loads or restraints

Additional loads or restraints should be selected and located in a way, that within the heated region additional compressive stresses arise. This can either be done by positioning of the element in a way that the dead load of the element leads to compressive stresses at the heated region. In some cases this load is not sufficient. An additional load can either be applied load controlled or path controlled.

Load control means that a defined weight is put e.g. in the middle of a beam that has been put onto two supports. During the whole heating and cooling cycle the weight will stay at its position when not lifted before by the workmen. The advantage of this procedure is that the additional load stays constant during the process which makes it easier for a prediction of the deformation. The disadvantage is that a load control can result in instabilities like the formation of a plastic hinge at the heating location when a

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large part of the section is heated or lateral torsional buckling of elements with small lateral stiffness. To avoid these instabilities the maximum bending moment should be restricted to 25% of the plastic moment resistance of the section.

Path controlled means that the deformation of the element is restricted. This can mean that a positive deformation during the heating cycle is obstructed or that an additional deformation in the desired direction is induced by mechanical means. Instabilities are less likely to occur, but the compressive stresses can hardly be calculated by simple means.

In both ways the heating patterns has the best effect when it is located in the vicinity of the load or restraint.

3 Heating patterns

All heating patterns have in common that the side of the component which shall be shortened is heated to induce the intended deformation. In the following the principal heating patterns to be applied on plate and beam structures are presented.

3.1 Types of heating patterns

3.1.1 Heat spots

A heat spot is, regarding the process, one of the simplest heating patterns. The torch is held on an indicated point for a defined time to heat up to the aspired temperature. By oscillation of the torch the diameter of the heat spot can be increased to the desired size. The resulting temperature field is concentric and has its maximum in the centre of the heat spot. During cooling the contracting forces take effect in all directions from the centre of the spot.

Heat spots are typically applied on plate structures with regard to deformations out of plane of the plate, in most cases for plate thicknesses smaller than 20 mm. Depending on the combination of nozzle and plate thickness the heating can either be superficial or through thickness. The term of “superficial” and “penetrative” heat was correlated to the cooling time between 500°C and 300°C, the so called t5/3, and in consequence to the penetrating action of the flame. In most cases a heat spot is a superficial heating pattern.

The main parameters of the heat spot are:

- Type of nozzle (concentrated or distributed heat flux distribution)

- Heat input of the nozzle q [J/s]

- Heating time th [s]

- Plate thickness tpl [mm]

Due to the static location of the torch the temperature rate decreases with time and the temperature field tends to a limit state of saturation. The maximum limit temperature of this limit state depends also on the combination of the nozzle (size, heat input) and the plate thickness. For this reason the correct selection of the nozzle is important: too small nozzles in combination with a large plate thickness will result in a low limit temperature and therefore very long heating times to induce the intended deformation, too large nozzles will have such a high heating rate which makes the process at less controllable.

The heating spot is normally used several times in different locations of a plate to reduce the deformation arisen from the fabrication process.

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Special forms of heating spots are lines of heat spots to reduce the heat input compared to a heat stripe or heat line.

3.1.2 Heat stripe, line heat

The heat stripe or line heat is a heating pattern which can be applied on both plate and beam components. The heating is performed by moving a torch along a line. During the longitudinal movement the torch can be additionally oscillated perpendicular to the moving direction to distribute the heat within a certain range. This is especially recommended for nozzles with a very concentrated flame to reduce the maximum temperature at the surface.

The process starts with heating up of the component at the starting point. After this region has reached the aspired temperature the movement of the torch is started with a defined velocity in the intended direction. After a certain time depending on the boundary conditions (size and heat input of nozzle, plate thickness) the temperature field reaches a quasi-stationary state. This quasi-stationary field travels along with the torch. Special attention should be given when reaching edges or corners of plates because due to the change in boundary conditions the temperature starts to rise near these locations, if the velocity is kept constant and, therefore, the velocity of the torch should be increased.

The contracting forces of a line heat act in transversal direction of the plate leading to a rotation angle at the heated location. But also in longitudinal direction contracting forces occur which can result in undesired deformations. Therefore special attention should be given to the supporting locations of elements to be straightened in transversal direction.

The main parameters of the line heat are:

- Type of nozzle (concentrated or distributed heat flux distribution)

- Heat input Q [J/s]

- Velocity v [mm/s] of the torch

- Plate thickness tpl [mm]

- Length of the line heat

- Oscillation parameters: amplitude and period of oscillation

The heat input can be related to the heated length by dividing the heat input q through the velocity which results in the heat input per unit length

qs = q / v.

A line heat is typically a superficial heating as the heating takes place from one side of the plate. The maximum temperature governs the minimum velocity of the torch which results in a thermal gradient over the thickness.

To reduce the effect of a line heat the process can be intermitted or even reduced to a line of heat spots. To enlarge the effect groups of line heats can be applied either subsequently or, with several torches, simultaneously.

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Beside plate structures line heats can also be applied on beams for out of square deformations.

These heating patterns have to be performed with special care because of the risk of a bending about one of the transversal axis. For the performance the beam should be placed onto the ground.

This effect can be utilized by placing a beam on two supports at the ends of the beam and by application of a longitudinal line heat.

3.1.3 Banded heat

A banded heat is a special form of the line heat in which the whole thickness of a plate is heated up (penetrative or through thickness heating). This can be done by heating form both sides of the plate or through repeated heating from one side.

Heat bands are mainly applied for beams to induce a deformation about one of the transversal axis.

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The main parameters of the line heat are:

- Type of nozzle (concentrated or distributed heat flux distribution)

- Heat input Q [J/s]

- Heating time [s]

- Plate thickness tpl [mm]

- Geometry of the band: width bband, length lband [mm]

- Oscillation parameters: amplitude and period of oscillation

The process for this heating pattern is given by an oscillating movement of the torch. For small bands the torch is oscillated along the band, while for wider bands the oscillation is perpendicular to the band direction and the torch is moved slowly along the band.

3.1.4 Vee heat

The vee heat is a special form of a heating pattern. The application lies within the deformation of beams about one of the transversal axis. Depending on the type of section the vee heat is a single or combined heating pattern:

For I- or H- sections about the strong axis it is a combination of a transversal heat band on the flange and a triangular heat band on the web, together forming a vee.

For T-sections it is either a transversal heat band or triangular band in the web depending on the direction of the bending.

For L-sections the application of a vee heat to induce a deformation about one axis leads, due to the eccentricity of the heating pattern, to a deformation about the other axis of the section.

It can also be applied on plates to induce a bending before the weld process.

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The main parameters of the vee heat are:

- Type of nozzle (concentrated or distributed heat flux distribution)

- Heat input Q [J/s]

- Heating time [s]

- Web and flange thickness tw, tf [mm]

- Geometry of the vee: height hband, width bvee or angle vee [mm] or [°]

- Oscillation parameters: amplitude and period of oscillation

According to the literature the heating of a vee shall be started at the apex of the triangle. With a gradually widening oscillation the torch is moved towards the wider end of the vee. For I or H sections about the strong axis the heating of the transversal band follows. The process is the same as for heat bands. From the tests performed within the project only a small influence could be seen from the sequence of heating proving this statement.

The following geometrical parameters of the vee are recommended:

Height: hvee = 0,5 – 0,67 hprofile

Width: bvee ≤ 120 mm

In general with larger angles of the vee the deformations increase due to the larger heated area. But for very large angles the probability of instabilities in the heated part increases especially when used together with large additional loads.

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4 Procedure

The procedure for a flame straightening process is divided in the following steps:

Preparation step:

- The material of the element has to be identified to choose the adequate upper temperature limit according to chapter 6. It should also be considered that for higher strengths the resulting deformations are lower.

- The structural element should be measured precisely. The deformations should be documented in a sketch or drawing.

- Depending on the distortion to be straightened or the deformation to be applied the type, size, location and sequence of the heating patterns has to be identified.

- The stiffness of the element has to be assessed and the necessity of additional weights or restraints has to be considered.

- If necessary, the type, size and location of additional restraints have to be selected.

- If available, the deformations may be estimated with the prediction means of chapter 7 of this guideline.

- Selection of heat input and/or heating time

- Selection of the nozzle according to the plate thickness

Process step:

- Adjustment of the flame

- Heating of the patterns

- Parallel or intermitted temperature control and measurement of the heating time

Cooling phase

128

5 Quantification and control of heat input

The quantification of the heat input is an important aspect for the certainty and the reproducibility of flame straightening procedures. Typically the used nozzles have a large range of gas flow to be applied. The adjustment of the torch depends on the experience of the operator.

For a clear adjustment of the torch the application of a device for the measurement of the gas flow is recommended. If these devices are not available the following procedures may help to quantify the heat input.

5.1 Quantification of effective heat input

The heat input by a gas nozzle has to be quantified by the lower heating value (LHV) of the gas. For the most common gases the values are:

LHVAcetylene = 57.120 J/l (48.700 kJ/kg)

LHVPropane = 93.000 J/l (46.300 kJ/kg)

The total heat input qu of a nozzle is calculated from the LHV by multiplication by the actual flow rate / of the nozzle

∙ /

The flow rate is typically given in l/min or m³/h. The typical flow rate of a nozzle is not a single value but a range of the possible flow rate (e.g. for a nozzle FB-A 9 the range of the acetylene flow rate is 3,12-4,15 m³/h). The actual value depends on the adjustment by the operator and should be measured.

To get the effective heat input q for the nozzle the total heat input has to be multiplied by the efficiency of the nozzle u. The effective heat input can be calculated from

∙

To quantify the energy input into the element the heat input q has to be put in relation to the heated element which can be done by division of the heat input by the velocity of the torch which results in a heat input per unit length:

For most nozzles the efficiency is not known, so that the following procedure maybe helpful to quantify the effective heat input respectively the value of efficiency.

5.2 Testing procedure to quantify the effective heat input

The following procedure allows an estimation of the effective heat input of the flame of an acetylene torch.

1. The following test equipment is required:

- A test specimen (The dimensions of the test specimen, the mass of water and the heating time depend strongly on the size of the nozzle. For nozzle n° 5 with a heat input of 4500 J/s a plate with dimensions 20x100x150 mm and a heating time of 30s resp. 60s was sufficient. For larger nozzles (nozzle FB-A 10 has an effective heat input of 17500 J/s) the dimensions of the plate shall be enlarged e.g. 20x200x300 mm)

- An acetylene torch

- An insulated container with a defined mass of water (As stated above the mass of water depends on the size of the torch. To have a measurable temperature increase of at least T = 5 K the mass of water can be assessed by the formula:

54187

tq

mwater .

Assuming a time period of 60s for a small nozzle n° 5 the mass of water has to be 12,9kg while for a large nozzle n° 10 it has to be 50kg.)

- At least two thermocouples to measure the temperature increase T of the water.

129

- To derive an efficiency of the nozzle u a gas flow meter is necessary to measure the flow rate of the acetylene and to quantify the total heat input.

2. The adjustment of the acetylene flame should be selected as for the flame straightening process.

3. With the torch the test specimen is heated up for a defined time period t (30s, 60s, etc.). During the heating the temperature of the test specimen should not exceed 600°C to exclude phase transformations and high amounts of evaporation of water.

4. Directly after heating the test specimen is cooled in the water. The temperature increase T of the water is measured by thermocouples. The temperatures have to be measured until there is no relevant change in temperature respectively the temperature of steel and water is nearly the same.

5. Afterwards the effective heat input can be calculated using the following equation:

01 TTmcTmcq waterwaterwaterwater [J/s]

with cwater = 4187 J/kg K. If the starting temperatures of steel plate and water are different, the temperature T0 has to be water calculated by:

steelsteelwaterwater

steelsteelsteelwaterwaterwater

cmcm

TcmTcmT

,0,00

6. By dividing the effective heat input q by the total heat input qu derived from the measurement with the gas flow meter the efficiency u can be calculated.

7. Measured values for different nozzles investigated within the project are given in Table A2-5.1.

Table A2-5.1 : heat input parameters for nozzles investigated within the project

Nozzle Type of nozzle Effective

heat input q

Total heat input

qu

Degree of efficiency

C2H2 gas flow

J/s J/s - l/h

FB-A 9 multi-flame rosebud 15.548 40.260 38,60% 3.018

FB-A 10 multi-flame rosebud 22.070 59.762 36,90% 4.480

N° 8 single flame 14.949 25.345 58,98% 1.900

Geyer A10 multi-flame rosebud 18.738 43.414 43,11% 3.255

7A single flame 11.550 28.941 39,94% 2.170

8A multi-flame rosebud 15.086 36.228 41,64% 2.716

5.3 Calibrating the heat input by numerical means

One possible strategy for calibrating the heat input is to perform some specific calibration flame tests. These tests basically consist of heating a calibration plate (preferable with thermal properties similar as those of the target material) instrumented with several thermocouples. In such a way, temperatures can be easily correlated with heating times or steel surface colours, and by means of numerical simulations, the total heat input can be estimated. This method for calibrating heat parameters can be a valuable tool for training workers in the workshop before flame straightening processes. Also the data obtained in this test can be used as the heat input parameters in a numerical simulation. As an example of this procedure, the details of two of these tests are summarised below.

In the two cases, the flame consisted of propane, with a mixture of 3 parts of oxygen and 1 part of propane and the torch was a type HARRYS 2290H4, with a capacity of 80000 to 158000 kcal/hour. Two steel plates with different geometric measurements were heated up. To do this, a different kind of flame was used in each plate:

1. Steel plate (1) (Figure A2-5.1). S355JR. Measurements: 570x470x50mm3. Static flame applied in the centre.

130

Figure A2-5.1: Steel plate (1) flame and thermocouples position

2. Steel plate (2) (Figure A2-5-2). S355JR. Measurements: 570x470x20mm3. Dynamic flame applied along the longest side.

Figure A2-5-2: Steel plate (2) flame and thermocouples position

In plate 1, drillings were made 10 mm below the top surface in five different positions, in the centre of the plate, and 20, 30, 50 and 70 mm from the centre (see Figure A2-5.1). To manage the temperature inside the steel of the plate, thermocouples were placed in these drill holes so the temperature evolution inside the steel, as a function of time and position, could be controlled.

In steel plate (2) three drillings were made, one in the centre of the plate, one 100 mm from the centre and one 200 mm from the centre and again these drillings were placed 10 mm below the top surface. A deformation transducer was placed to verify the movements in the centre of the plate. This distribution was as shown in Figure A2-5-2.

The operators try to heat up the steel to a particular temperature based on the colour of the steel (see this appendix, part 2 “Equipment, visual control of temperature”) and on the time the flame has been applied over the plates. The operators try not to go beyond 650 ºC. In Figure A2-5.3 the results of the thermocouples in plate 2 are shown.

This test also achieved the calibration of the flame for the heat input in the finite elements simulation.

With the data provided by the thermocouples and transducer in the first stage, the second step of modelling the flame was initiated. The flame model was set as a Gaussian heat distribution applied to the material. It was decided to implement the 2D model of the flame, instead of a 3D model of the flame, as this was the optimum way to represent the flame in the ANSYS code. Therefore, the heat flow density of the flame follows the expression:

2

05,0

2

2

3

205,0

3 r

r

rk er

qe

kqrq

Where q̂ is the maximum heat flow density and k is the concentration factor of the Gaussian Distribution.

205.0

3

rk

131

Figure A2-5.3: Temperature results in plate 2

Using the data collected in the laboratory tests, real temperature-time graphs were represented and the parameters q̂ and k were adjusted by means of several numerical iterations, obtaining at the end temperature-time curves and comparing them with the real ones.

First the parameter q̂ is fixed and then the parameter k varies its value until getting the best fit. Then,

when parameter k is obtained, the parameter q̂ varies its value until getting also the best fit. Figure A2-5.4 show how this process is performed.

In Figure A2-5.4 (left) the parameter r0.05, which k is function of, is fixed at a value of 85 mm, based on previous experience, and the other parameter, q̂ , vary its value to find a curve that fit with the real one.

The chosen value for q̂ was 0.55*5.21 J/(mm2.s) (55% of Rykalin parameter). In the Figure A2-5.4

(right) the parameter q̂ nonw is fixed at 0.55*5.21 J/(mm2.s), as it was defined in Figure A2-5.4, and r0.05 varies its value until a curve fit with the real one. A value of 85 mm was chosen for parameter r0.05 and the final curve is modelled with the parameters 0.55*5.21 J/(mm2.s) and 85 mm.

Figure A2-5.4: Curves for adjusting flame parameters

Using the coefficient of the heat flux distribution the effective heat input can be calculated to

sJrq

qr

qq /21680

3

8521,555,0

3

ˆ3ˆ

2205,0

205,0

Flame ^q,r0.005=85

0

100

200

300

400

500

600

700

800

900

0 100 200 300 400 500 600 700 800 900 1000

Time (s)

Te

mp

era

ture

ºC

Real test results

ANSYS curve,^q=50, r=85

ANSYS curve,^q=55, r=85

ANSYS curve,^q=45, r=85

ANSYS curve,^q=60, r=85

Flame ^q=0.55, r00.5

0

100

200

300

400

500

600

700

800

900

0 200 400 600 800 1000 1200 1400 1600 1800 2000

Time (s)

Te

mp

era

ture

ºC

Real test results

ANSYS curve: ^q=55, r=70mm

ANSYS curve: ^q=55, r=85mm

ANSYS curve: ^q=55, r=100mm

132

6 Avoiding an impact on the material

6.1 Temperature range for flame straightening

The influence of the flame straightening on the mechanical properties of the plate depends on the flame straightening temperature relative to the initial heat treatment (delivery condition) of the plate / beam. The flame straightening has to be applied within a defined temperature range. The lower limit of that range is set by insufficient straightening; the upper limit has to be respected in order to exclude a deterioration of the mechanical properties depending on the delivery condition.

6.2 Delivery conditions

Plates and sections are available in different delivery conditions: as rolled (AR), normalised (N), quenched and tempered (Q) and thermomechanically rolled (M). In Figure 6.1 the temperature-time diagrams of these conditions are shown.

6.2.1 As rolled (AR)

After heating at temperatures of about 1100 °C the rolling of the slab takes place in the austenitic state, a crystal structure that is stable at high temperatures. After that the plate cools on calm air and the “as rolled” condition (AR) is achieved (Process A in Figure 6.1).

Figure A2-6.1: Delivery conditions

6.2.2 Normalised (N)

To get a more homogenous microstructure an additional heat treatment can be performed. The plate is reheated just above the ferrite-austenite transformation temperature (about 800 – 900 °C, depending on the carbon content) and is cooled on calm air again. By this treatment the steel transforms from ferrite and pearlite to austenite and back again. This leads to a refined microstructure of ferrite and pearlite, which is called the normalised condition (N). By normalising steel grades with moderate strength and toughness requirements up to S460N can be produced (Process A + B in Figure 6.1).

6.2.3 Quenched and tempered (Q)

The temperature-time-diagram for the quenching and tempering process is quite similar to the one for normalising (Process A + C in Figure 6.1). After hot rolling and cooling the plate is reheated above the transformation temperature, so that carbon can dissolve in austenite, but then cooling is not performed on calm air, but in water (quenching) or in another medium that cools fast enough, so that there is no time for the formation of ferrite and pearlite which needs a diffusion process. The carbon stays dissolved and at room temperature the microstructure mainly consists of martensite, a distorted structure that has a high strength but a low toughness. With an additional tempering process the crystal latter has the possibility to relax with the effect that strength decreases while toughness increases, so that a material with a satisfactory combination of tensile and toughness properties can be produced. The effect of the tempering on the mechanical properties is shown in Figure 6.2.

133

Figure A2-6.2: Influence of increasing tempering temperatures on the tensile properties (left) and on the Charpy V

transition temperature (right) - S890QL, 60 mm

High strength quenched and tempered steels need particular care when thermal straightening is applied.

- During heat treatment the plate was tempered at 600°C. To avoid a local softening by the "heat treatment" during the straightening process the peak temperature at the hot spot shall clearly be kept below this initial tempering temperature.

- Consequently 600°C shall generally be respected as the upper limit for the straightening process. Only if a limited local softening of the surface area is acceptable (as it will be for several applications) somewhat higher peak temperatures may be suitable. In any case the maximum temperature shall be clearly below 700°C because a partial transformation to austenite must be avoided. Partial or full transformation to austenite can cause local embrittlement and/or hardening.

- For the heat straightening, if the tempering temperature is lower than 600°C, it is recommended to respect the limit of the tempering temperature - 30K.

- Preliminary tests on the same steel type and microstructural investigations and mechanical tests may be required to prove the suitability of the intended processing conditions.

6.2.4 Thermomechanically rolled (M)

Another way to get steel with high strength is to create a microstructure with an extremely fine grain. The smaller the grain size is the higher are the tensile and toughness properties. The thermomechanical rolling (TM or TMCP) is a method to realise such a fine grained microstructure by a skilled combination of rolling steps at particular temperatures and a close temperature control (Processes D to G in Figure 6.1). The gain in strength obtained by the grain refinement allows reducing effectively the carbon and alloying content of the TM-steel compared to normalised steel of the same grade. The improved weldability that results from the leaner steel composition is a major advantage of TM-plates. The applied "rolling schedule" is individually designed, depending on the chemical composition, the plate thickness and the required strength and toughness properties. Some typical TM-processes are shown in Figure 6.1. Especially for thick plates an accelerated cooling (ACC) after the final rolling pass is beneficial for the achievement of the most suitable microstructure as it forces the transformation of the elongated austenite grains before recrystallisation can happen. For very thick plates and high strength steel grades a tempering process can be used after accelerated cooling.

6.2.5 Microstructure

Figure 6.3 shows microstructures of the different delivery conditions. It is easy to identify the typical microstructure of normalised steel which is dominated by ferrite and pearlite. A direct comparison with TMCP structures shows two main differences. First, there are less black areas, a result of the lower carbon content and second, the smaller grain size, which is smallest when ACC is performed. A completely different appearance has the quenched and tempered steel. The martensite that is formed by displacive transformation shows a acicular (needle-shaped) microstructure.

134

Figure A2-6.3: Microstructures of various delivery conditions

6.3 Upper limit of applicable temperature

The upper limit of the temperature range shall not be exceeded in order to retain the mechanical properties of the plate after subsequent cooling.

Two different flame straightening procedures have to be distinguished because they require different temperature limits: superficial heating and full section heating.

The following maximum applicable temperatures shall apply to steels acc. EN 10025:

Delivery Condition Upper limit of applicable temperature 5) 6)

Superficial heating 1) Full section heating 2)

As rolled (AR)

900 °C 700 °C Normalised (N)

Thermomechanically rolled (M)

up to S460 4)

Quenched and tempered (Q)

up to S690 800 °C 3)

30 °C below tempering temperature

of the product

1) In this case first case the plate is only superficially heated. The heat input related to the plate thickness is small. Due to the steep temperature gradients in the through thickness direction cooling is very fast. Cooling speed is in the range of high heat input welding (example: line heating).

2) In this case the entire cross section of the plate is heated. The heat input related to the plate thickness is high. The heated area cools down slower than in the first case. The larger the heated area the slower the cooling. Air cooling represents the very extreme of such procedure (examples: wedge heating, triangular heating). In case the full section heating is elongated to longer holding times the maximum applicable temperatures should be reduced by 50 K for all steels except for Q-steels where no further reduction is necessary (see [3]).

3) In the area of the plate or beam where the tempering temperature is exceeded a local reduction of the mechanical properties may occur. In consideration of the plate or beam as a whole this can be neglected, depending on the application. In cases where the local reduction cannot be neglected, the maximum temperature should be limited to 30 °C below tempering temperature of the product

4) Thermomechanically rolled steels higher than S460 are not covered by EN 10025 and were not investigated in the project. In case of these steels (S500 to S690) acc. to [3] and [57] for superficial heating maximum temperatures up to 900° C and for full section heating maximum temperatures up to 600° C are applicable.

5) Adequate measurement devices are given in chapter 2.2.1 of this guideline. Procedures for the measurement of the temperature are given in chapter 7 of this guideline. If the control of the maximum temperatures is not applicable or uncertainties exist about the correct temperature measurement procedure tests should be performed to ensure the original mechanical properties of the material.

6) In case of uncertainty please ask the producer of the steel product.

135

7 Prediction means and examples

The success of a flame straightening process is given when two conditions are fulfilled.

- No detrimental effect has taken place within the material resp. the mechanical properties of the material have not changed due to the heating process. This is supposed to be the case when the maximum allowable temperatures given in chapter 6.3 are not exceeded.

- The aspired bending has been applied or the distortion from prior fabrication processes has been straightened.

The success of the straightening process depends on the manual adjustment and handling of the torch. To achieve a good prediction of the results of a flame straightening process the control of temperature, heat input and heating time is of major concern as these are the input parameters for the estimation of the deformation.

The control of the heat input should be done by precise adjustment of the torch. For the applied torch a quantification of the heat input should be done according to chapter 5 of this guideline in dependency of the gas flow beforehand.

Devices for the control of the temperature are given in chapter 2.2.1 of this guideline. For the control of the temperature several cases have to be distinguished:

For superficial heating patterns the maximum temperature in the heating pattern is of concern. Directly after removal of the torch the temperature should be measured. The time gap between removal of the torch and measurement should be minimized (not more than 5s after). Depending on the plate thickness the cooling rates are larger than 10K/s, so that to the measured temperature at least 30-50 K should be added.

For penetrative or full section heating patterns and for thin plates (< 20 mm) the temperature should be measured on the surface, which is opposite to the heating surface, while heating or, if not applicable, on the heated surface not more than 10 s after the flame was removed. For penetrative or full section heating patterns and for thick plates (> 20 mm) the temperature should be measured on the heating surface under the torch approximately 10 s after the flame was removed. This procedure is based on temperature measurements in the trials and in the numerical simulations. The following figure shows the locations for the temperature of control of a vee heat:

For other heating patterns the locations should be selected analogously. If these procedures are not applicable during the heating process of the element the correct process parameters should be verified in prior procedure tests.

In the following prediction means for line heats on plate structures and heating patterns on beams for bending about the weak and the strong axis are presented which were derived on the basis of the experimental, numerical and analytical investigations.

7.1 Application of prediction means for plates with line heats

7.1.1 Procedure for the prediction of the rotation angle

The prediction of the rotation angle of the line heat is based on the experimental and numerical investigations of the project. It depends on the main parameters, which are

- Type of nozzle (concentrated or distributed heat flux distribution)

136

- Heat input Q [J/s]

- Velocity v [mm/s] of the torch

- Plate thickness t [mm]

For the prediction the heat input q has to be related to the heated length by dividing the heat input q through the velocity torch which results in the heat input per unit length

qs = q / v.

With this input parameter the maximum temperature can be estimated from Figure A2-7.1.

Figure A2-7.1: maximum temperature vs. heat input per unit length

To derive the estimated deformation the heat input has to be related to the plate thickness t

∙

Using this parameter the rotation angle can be derived form Figure A2-7.2 by multiplication of the value taken from the graph by m and by division by the plate thickness t. The parameter m depends on the steel grade and, for steel grades S235 to S460, can be calculated by

4,58 ∙ 10 4,76 ∙ 10 ∙ , fy in [MPa]

For steels S690 and S890 only lower values of the heat input per unit length and thickness up to 200 J/mm² fit to the estimation. Up to this value m = 0,0007 rad mm³/J fits best. All values of m are given in Table A2-7.1.

Table A2-7.1 : Slope m for different steel grades (average values)

Steel grade S235 S355 S460 S690 S890

Slope [rad mm³/J] 0,0033 0,0028 0,0022 0,0007 0,0007

137

Figure A2-7.2: thickness x rotation angle / m vs. heat input per unit length and thickness

7.1.2 Example

For the presentation of the prediction procedure a plate with a thickness t = 20 mm is considered with T-stiffeners which have a distance of 1000 mm each. The plate has a width of 1500mm and a length 4500mm. The material consists of a steel S235JR. Due to the welding process of the stiffeners to the plate distortions on form of angles have been observed.

A flame sraightening process in form of line heats is considered.

Figure A2-7.3: Flame straightening process of a plate with T-stiffeners (left: performance, right: location of line heats (FS1 to FS 7)

A multi-flame rosebud nozzle 8A according to Table A2-5.1 has been selected for the flame straightening process. The gas flow is considered to be 2700 l/h. Heat lines FS3, FS 5 and FS 6 shall be predicted. The measured gas flow was 2600 l/h for FS5 and FS6 and 2400 l/h for FS 3.

The prediction for the three line heats is given in Table A2-7.2.

Table A2-7.2 : Prediction of the deformation of the considered line heats

Line heat

Time s

Length cm

v mm/s

q J/s

qs J/mm

qs / t J/mm²

t / m -

rad

pred.

mm meas.

mmFS3 383 150 3,9 13285 3392 170 124 0,021 10,3 11,0

FS5 383 150 3,9 14417 3681 184 140 0,023 11,6 12,0

FS6 298 150 5,0 14417 2864 143 96 0,016 7,9 8,3

From the records of the deformation transducers during the test the deformations can be read. The deformation transducers are located in distance of 500 mm on both sides of the line heat (see Figure A2-7.3, right). The results are shown in Figure A2-7.4 and Figure A2-7.5.

138

Figure A2-7.4: Deformation record for line heat FS3

Figure A2-7.5: Deformation record for line heats FS5 and FS6

A comparison of the results with the prediction is given in Table A2-7.2. From Figure A2-7.1 the maximum temperatures have been estimated to 500° C for FS6 and 600° C for FS3 and FS5 (t = 20 mm, FB-A 9). The Records of the thermo vision camera prove that the prediction of these temperatures is good.

4-6 =

5,5 mm

1-3 =

5,5 mm

4-6 =

7,6 -9,1 mm

1-3 =

11,3-12,7 mm

139

Figure A2-7.6: Records of the thermo vision camera ( = 0,85) of the line heats (left: FS3, middle: FS5, right: FS6)

7.2 Application of analytical model 1 to I and H profiles

The section provides the application of the analytical method 1 presented in 4.1.1. The effect of V heat on strong and weak axis bands on I and H profiles in relation to the level of external load is given.

The method is based on the preposition of uniform temperature field. The mechanical properties of structural steel at elevated temperature are the basis for the calculation of the critical temperature. Figure A2-7.7 presents normalized proportional stress in relation to temperature, gathered from different sources. Eurocode curve (EN 1993-1-2) and S235; lower bound present lower bound while the other curves present the experimental results provided by Chen et al. [64]and RFCS project OPUS (for S235; mean values).

Figure A2-7.7: Reduction factor for the proportional stress at elevated temperature

The analytical method 1 presented in 4.1.1 may be rewritten in more compact form. If the equation that presents equilibrium at the first critical temperature

,1, crit

veep T p

c T

Mf f

A c , (1)

is divided by Mpl and the plastic moment is replaced by fy Wpl, it follows:

p pl vee

y c T pl

f W M

f A c M . (2)

The left hand side of the previous equation presents the reduction factor for the proportional stress, while the right hand side of the equation presents the product of the geometrical factor kg and load factor kf. The normalized for of Eq. (1) follows:

0

0.2

0.4

0.6

0.8

1

0 200 400 600 800 1000 1200Temperature [°C]

kp=f

p/f

y(20

°C)

High strength steel;Ju Chen et al.

Mild steel; Ju Chen et al.

S235; mean values

S235; low er bound

EN 1993-1-2

140

; ; ;p pl veep g f p g f

y c T pl

f W Mk k k k k k

f A c M . (3)

It should be considered that the plastic bending resistance should be calculated with actual value of the yield stress to obtain realistic load factor.

7.2.1 Strong axis bands

The V heats are usually applied from 2/3 up to full depth of the cross-section (bottom flange is unheated at full depth). It has to be considered that the unheated portion of the cross-section increases the stiffness of the beam which is not accounted for in the presented method. Therefore the deformation of V heat is presented only for V heats of full depth. Figure A2-7.8 presents the geometrical factor kg for I and H profiles for full depth of V heat. Due to similar shape of the profiles the factors lay within a relatively small interval between 0.75 and 0.9 for all sizes of IPE, HEA, HEB or HEM profiles.

Figure A2-7.8: Geometrical factor kg for IPE, HEA, HEB, HEM profiles for full depth of V heat

The following figures illustrate the rotation of the V heat in relation to the load factor kf for full depth of V heat. The curves present the initial angle of V heat from 10°to 50°. The critical temperature depends only from the load factor and is the key parameter in the presented figures. The rotation of the heat is an informative parameter. If the temperature is higher than the critical temperature, the rotation may drastically increase and if it is considerably lower, only a fraction of the rotation may be expected. The rotation of the V heat is based on the deformation due to the thermal strain as presented in the description of this analytical model. Therefore higher temperatures result in higher rotation. In Figures A2-7.9 and A2-7.10 the critical temperatures are calculated for two material models presented which are defined in Figure A2-7.7. The material model for S235 is based on the experimental values of the proportional stress at elevated temperatures, while the Eurocode model presents a lower bound of the proportional stress. Therefore, the temperatures are considerably lower in Figure A2-7.10 than in Figure A2-7.9.

0.75

0.80

0.85

0.90

0 200 400 600 800 1000Cross-section

k g

HEM

HEB

HEA

IPE

141

Figure A2-7.9: Rotation of the V heat at the critical temperature for IPE 400; material model S235 (mean values); full depth of V heat

Figure A2-7.10: Rotation of the V heat at the critical temperature for IPE 400; material model EN 1993-1-2; full depth of V heat

The influence of the geometrical factor is presented in Figure A2-7.11, where the critical temperatures are calculated for IPE 400 (kg = 0.79) and HEA 400 (kg = 0.87) profiles. It may be concluded that the influence of the geometrical factor is for IPE, HEA, HEB or HEM profiles has a small effect on the critical temperature.

Figure A2-7.11: Rotation of the V heat at the critical temperature for angle of V heat 30° and S235 (mean values) ; full depth of V heat

687

687

687

687

687

671

671

671

671

671

745

745

745

745

745

706

706

706

706

706

784

784

784

784

784

859

859

859

859

859

967

967

967

967

967

1083

1083

1083

1083

1083

0

2

4

6

8

10

12

0 0.1 0.2 0.3 0.4Load factor k f

Rot

atio

n o

f th

e V

hsa

t[m

rad

]

10 20 30 40 50

V angle [°]

546

546

546

546546

523

523

523

523523

590

590

590

590590

568

568

568

568568

620

620

620

620

620

658

658

658

658

658

696

696

696

696

696

882

882

882

882

882

0

2

4

6

8

10

0 0.1 0.2 0.3 0.4 0.5

Load factor k f

Rot

atio

n of

the

V h

eat

[mra

d]

10 20 30 40 50

V angle [°]

745

859

967

1083

784

687 671706727

694

770

831

1072

946

677 6603.5

4

4.5

5

5.5

6

6.5

7

0 0.05 0.1 0.15 0.2 0.25 0.3 0.35 0.4Load factor k f

Rot

atio

n o

f th

e V

[m

rad

]

IPE400

HEA400

142

7.2.2 Weak axis bands – V heat

The calculation of the critical temperature for V heats applied in the weak axis of I or H profile is similar to the strong axis orientation. The special case presents the full depth of V heat. In this case there is no unheated part of the cross-section. At the critical temperature the cross-section fully yields at fp(Tcrit) and the neutral axis forms accordingly. Therefore the geometrical factor is equal to 1 and the load factor is equal to the reduction factor for the proportional stress.

; ; 1;p veep f p g f

y pl

f Mk k k k k

f M . (4)

7.3 Application of numerically derived prediction means for beams

The following predictions are based on the study of the results obtained using finite elements simulations by ANSYS and validated by large scale laboratory tests. The finite elements simulations were performed on short beams, 50 cm long, (see point 3. Numerical simulations, 3.3 bar shaped structures for further details) but these results can be extrapolated to an entire beam by applying the proposed method. In each one of the figures of this section are represented the values of rotation for one single heated area. These values are directly related with the final deflection of the profile. In the parametric study summarised here, the impact on the deflection of the beam of the next four different parameters have been studied: profile type (HEA 300, IPE 450 and HEM 340), material grade (S235, S355, and S460), level of additional restraints (applied forces) and temperature-time of heating.

7.3.1 Influence of temperature

The temperature reached as a consequence of the time the flame is applied over a steel structure is the most decisive factor in a thermal process.

The rotation angle in the heated area of a beam, as a function of temperature for the reference cases, is represented in Figure A2-7.12, Figure A2-7.13 and Figure A2-7.14 for three different profiles. Three curves are shown in each graph representing three different levels of additional loads (0, 12.5 and 25% of the plastic moment for the profile).

Figure A2-7.12: HEA 300 temperature-rotation angle graph

HEA 300

0.00E+00

1.00E-03

2.00E-03

3.00E-03

4.00E-03

5.00E-03

6.00E-03

7.00E-03

0 100 200 300 400 500 600 700 800 900 1000 1100

Temp

Ro

tati

on

(ra

dia

ns

)

0Mp0.125Mp0.25Mp

143

Figure A2-7.13: IPE 450 temperature-rotation angle graph

Figure A2-7.14: HEM 340 temperature-rotation angle graph

As can be observed in these graphs (Figure A2-7.12, Figure A2-7.13 and Figure A2-7.14), no deflection, apart from the mechanical one, can be expected below a certain level of temperature. After that level, an approximate linear tendency is followed by the relationship between rotation and temperature.

7.3.2 Influence of bending moment

The figures below show the rotation angle as a function of the bending moment for three steel profiles commonly used in construction.

Figure A2-7.15: HEA 300 bending moment- rotation angle graph

IPE 450

0.00E+00

5.00E-04

1.00E-03

1.50E-03

2.00E-03

2.50E-03

3.00E-03

3.50E-03

0 100 200 300 400 500 600 700 800 900 1000 1100 1200

Temp ºC

Ro

tati

on

(ra

dia

ns

)

IPE 450 0.25MpIPE 450 0.125MpIPE 450 0 Mp

HEM 340

0.00E+00

2.00E-03

4.00E-03

6.00E-03

8.00E-03

1.00E-02

1.20E-02

1.40E-02

0 100 200 300 400 500 600 700 800 900 1000Temp

Ro

tati

on

(ra

dia

ns

)

0*Mp0.125*Mp0.25*Mp

HEA 300

0.0E+00

1.0E-03

2.0E-03

3.0E-03

4.0E-03

5.0E-03

6.0E-03

7.0E-03

0 20000 40000 60000 80000 100000 120000 140000

Bending Moment N*m

Ro

tati

on

(ra

dia

ns)

HEA 300 1000ºC

HEA 300 850ºC

HEA 300 760ºC

0.25Mp

0.125Mp

144

Figure A2-7.16: IPE 450 bending moment- rotation angle graph

Figure A2-7.17: HEM 340 bending moment- rotation angle graph

Obviously, as it can be observed in Figure A2-7.15, Figure A2-7.16 and Figure A2-7.17, higher values of bending moment cause higher values of deflection in the beam. The relationship between rotation-deflection and bending moment, for the values of bending moment used in this study, is proportional to a straight line until an inflection point is reached when this proportionality stops being rectilinear.

7.1.3 Influence of material grade Figure A2-7.18 represents the rotation angle as a function of the steel grade for a given value of bending moment. The profile studied is a HEA 300.

Figure A2-7.18: Rotation as a function of the steel grade

IPE 450

0.0E+00

5.0E-04

1.0E-03

1.5E-03

2.0E-03

2.5E-03

3.0E-03

3.5E-03

4.0E-03

0 20000 40000 60000 80000 100000 120000 140000 160000

Bending Moment N*m

Ro

tati

on

(ra

dia

ns

)

IPE 450 1000ºC

IPE 450 870º

IPE 450 700º

0.25Mp

0.125M

HEM 340

0.0E+00

2.0E-03

4.0E-03

6.0E-03

8.0E-03

1.0E-02

1.2E-02

1.4E-02

1.6E-02

1.8E-02

2.0E-02

0 50000 100000 150000 200000 250000 300000 350000 400000 450000

Bending Moment N*m

Ro

tati

on

(ra

dia

ns)

HEM 340 760ºC

HEM 340 950ºC

HEM 340 840ºC

0.25Mp

0.125Mp

rotation angle - steel grade

0.00E+00

1.00E-03

2.00E-03

3.00E-03

4.00E-03

5.00E-03

6.00E-03

7.00E-03

Steel grade

Ro

tati

on

an

gle

(ra

dia

ns

)

Mb = 0Mb = 20000Mb = 55000Mb = 100000

S235 S355 S460

145

The main conclusion for the parametric study of the material grade, is that, as it was expected, higher rotation angles are obtained for lower yield strength materials.

7.3.3 Example 1: Vee heat about the strong axis

Profile: HEA 300, 6 meters beam Steel grade: S355J2

The test consists in a 6 meters beam bending along its strong axis with two supports at its edges. The heat was applied in two steps, the first step with an additional load (F) of 70 kN in the middle of the beam and the second step with no additional load.

The setup of the test can be seen in Figure A2-7.19 and Figure A2-7.20. Thermocouples T1 and T2 were placed in the vee 9, as it can be observed in Figure A2-7.21 and a laser deformation transducer was placed in the middle span. The heating pattern consisted of a vee shape on one side of the web and a band in the flange. The dimensions are summarised in the Figure A2-7.21. The results of the test, -temperature and displacement records-, can be observed in Figure A2-7.22 and are summarised in Table A2-7.3.

Figure A2-7.19: Test setup. The first step in red, the second in blue

Figure A2-7.20: Flame straightening process of the profile.

Figure A2-7.21: Test setup. Thermocouples and vees and band shape

146

Figure A2-7.22: Temperature and deflection over time

Step Thermocouple Max. Temperature Deflection

1 T1 902ºC

29.85 T2 862ºC

2 - -

11.62 - -

Final - - 41.47 Table A2-7.3: Summary results

7.3.4 Example 1: Vee heat about the weak axis

Profile: HEA 300, 6 meters beam Steel grade: S355JR

The test consists of a 6 meters beam bending along its weak axis with two supports at its edges (Figure A2-7.23, Figure A2-7.24). The heat was applied in one step with an additional load (F) of 30 kN in the middle of the beam.

The position of the thermocouples as well as the shape of the V heat can be observed in Figure A2-7.25. The laser deformation transducer is the same as in Test 10.1 as is the shape of the V. The results of the test, -temperature and displacement records-, can be observed in Figure A2-7.26 and are summarised in Table A2-7.4.

Figure A2-7.23: Test setup.

147

Figure A2-7.24: Flame straightening process of the profile

Figure A2-7.25: Thermocouples position and shape of the V

Figure A2-7.26: Temperature and deflection over time

Step Thermocouple Max. Temperature Deflection

1 T1 967ºC

23.3 mm T2 961ºC

Table A2-7.4: Summary of results

7.4 Further prediction means for beams (model 2)

An additional prediction mean (model 2) is available which was derived analytically to calculate directly the deformations for beam structures with continuous and discontinuous heating patterns. By model 2 the plastic strains are derived using a fibre model of the section at different locations of the heating patterns. Assuming the hypothesis of Bernoulli and linear elastic – ideal plastic material behaviour the strain distribution is calculated in this way that the resulting stress distribution fulfils the equilibrium with the internal forces at the location of the heating pattern. The background of this model is given in chapter 4.1.2 of the final report.

By this model the dependencies of different parameters of heating pattern can be investigated and the optimum parameters for the heating process can be found. For example, in Figure A2-7.27 and Figure A2-7.28 the moment-rotation curves for a vee heat process on an IPE 450 profile, steel grade S235 with a vee heat about the strong axis (hvee = 300 mm, bvee = 80 mm) are given for different steel grades and various heights of the vee.

148

The fibre model of different sections including the described procedure has been established in an Excel-Worksheet and will be available for download on the project web page (link, see Appendix 1) parallel to the publication of the final report.

Figure A2-7.27: strain and stress distributions (IPE 450 profile, vee heat: 300 mm/80 mm)

Figure A2-7.28: strain and stress distributions (IPE 450 profile, S235, vee heat: width 80 mm)

8 Quantification of displacement

Before and after the flame straightening process the structural element has to be measured to quantify the deformations. In practice, the easiest and cheapest way to measure a deformation is to use a rope or a browning rod. For more complex structures or more precise results the use of a levelling device is recommended.

For the curvature along the weak and the strong axis, this concept could be easily applied. A rope should be tightened between the two extremities where the bending should be measured. It is advised to measure the bending at several locations.

In order to generate a bending along the strong axis, the section is set in the “I” shape position. In order to measure the curvature induced by the flame straightening, it is advised to rotate the section of 90° and dispose it in “H” shape (Figure A2-8.1A).

To induce a curvature along the weak axis, the section is set in the “H” shape position. To measure the bending, it is recommended to rotate the section of 90° (to the “I” shape) and perform the measurement as indicated on the Figure A2-8.1 B.

A – Bending along the strong axis B – Bending along the weak axis

Figure A2-8.1 : Principe to measure the bending with a rope tightened between the two extremities

Others kind of deformations could also be detected on sections, as out-of-square defect. The out-of-square could be constant on the length or the section or not. This could affect only one flange or the two flanges and not be identical for the two flanges. The out-of-square can be estimated by the difference between the height of the flanges 1-3 and flanges 2-4, as illustrated in Figure A2-8. 2A. The out-of-square is the sum of the two distance identified as “k and k’ ”. As the flange 1-2 is straight, for instance,

149

the total out-of-square is identified for the flange 3-4. If the out-of square is not identical for the two sides, the use of a square could be necessary to identify the out-of-square for each flange. The picture below represents a square used for one of the trials (Figure A2-8. 2 B) and the principle (Figure A2-8. 2 C).

A) Definition of the out-of-square

B) square used in a workshop to measure the out-of-square

C) principle of a square to measure the out-of-square

Figure A2-8. 2: Definition and measure of the out-of-square defect

150

Appendix 3: Guideline for the numerical simulation of flame straightening

1 Material model

The material model should include the adequate material properties which are dependent on the temperature. For the thermal calculation density, conductivity and specific heat are required. For the thermo-mechanical calculation the Young’s Modulus, Poisson’s ratio and thermal expansion coefficient should be available as well as flow curves of the material. A kinematic hardening law should be used.

Material properties can be found in various references: [64][66][67][68][69][70][71][72][73][74][75].

2 Numerical model of flame

The heat transfer from the flame to the surface of the element is a type of a forced convection. By this way the flame generates a heat flux on the surface of the element. Therefore, the flame is represented by surface heat flux. It was shown that the temperature field is greatly dependant on the heat flux density. Therefore a Gaussian model of flame should be accounted for. For detailed information see chapter 3.1.1 of the final report.

In most cases the flame moves along the surface. The path and velocity should also be considered. Such presentation captures also the temperature peaks on the surface. In certain finite element software, like Abaqus, the definition of moving heat flux is not supported. In such case the heat flux may be defined through a contact with thermal properties (radiation, convection, conduction) on a separate body to which the path and velocity may be given.

3 Finite element selection

For thermo-mechanical analyses the finite element selection is limited to the ones that have displacements and temperature as active degrees of freedom in all nodes. Beside solid element, also shell elements may be used. Appropriate integration scheme (Simpson's integration rule) should be defined if the surface temperature is required. Appropriate mesh density is required, especially for the analysis of the stress field.

4 Initial residual stress

The residual stress due to element fabrication (flame cutting, rolling, welding...) should be introduced to the numerical model to obtain more realistic response of the structure.

5 Type of the analysis

There are several procedures for heat transfer and thermal-stress analysis. Heat transfer analysis calculates the temperature field without knowledge of the stress/deformation state. Pure heat transfer problems in general can be transient or steady-state and linear or nonlinear. If the stress/displacement solution is dependent on a temperature field but there is no inverse dependency, a sequentially coupled thermal-stress analysis can be conducted. Sequentially coupled thermal-stress analysis is performed by first solving the pure heat transfer problem, then reading the temperature solution into a stress analysis as a predefined field. A coupled temperature-displacement procedure is used to solve simultaneously for the stress/displacement and the temperature fields. A coupled analysis is used when the thermal and mechanical solutions affect each other strongly. In general in flame straightening simulations the thermal results affect the mechanical results, but the mechanical results do not affect the thermal simulation. Therefore a stepwise simulation of the process is sufficient. Further instructions regarding the application of the analysis procedures are given in the software's manual (e.g. Abaqus Analysis User's Manial 6.10).

6 Stepwise guiding procedure for the numerical simulation

This guideline briefly explains the methodology for performing a numerical simulation of a bar-shaped structure using the ANSYS software code.

The method uses a non-coupled analysis. In a first instance, a thermal analysis has to be performed, the thermal loads then being transferred to a structural analysis. An optimum simulation process can be carried out in a reasonable time and using a conventional computer. A complete beam can be simulated in only one day with a conventional 8 GB RAM computer. The thermal simulations for one part of the beam can last 2 hours and the structural part takes 3 hours.

151

Step 1: First of all, the heating pattern, the number of heated areas and the additional restraints, if necessary, have to be defined. For more information go to Appendix 2 or chapter 3 of the final report.

Step 2: The profile has to be divided into equal short beams corresponding to one single heating area (see Figure A3-6.1). The external bending moment applied to each short beam as a consequence of the application of the additional restraints has to be calculated. The total deflection of the beam will be estimated by integrating the bending angles of each one of the short beams.

Figure A3-6.1: Simplification of the entire beam in short beams

Step 3: From a geometrical point of view, the only entity that has to be modelled is one of the short beams. As the heating details for all the short beams are usually the same, the thermal part of the simulation only has to be performed once. Using the results of the thermal simulation as inputs for the mechanical model, in theory, one mechanical simulation should be carried out for each short beam taking into account that the different bending moments applied in the different sections of the profile. However, as outlined in Section 3, with 3 single simulations, the trend for rotations versus bending moments can be defined.

For thermal simulation, it is recommended the use of the element type PLANE55, which is a 2D four node element, for defining the profile and then extruding it into elements type SOLID70 which is an eight-node element with conductivity capability. For the structural analysis an order called ETCHG, TTS is used to transform the thermal elements PLANE55 into a structural elements PLANE42, which is also a four-node element.

The beam geometry is simple, so the meshing can consist of quadrilateral elements lined-up with the longitudinal axis of the beam. In order to provide the beam with a higher element density in the area to be heated, smaller elements was used in it (see Figure A3-6.2).

Figure A3-6.2: Meshed Model of short beam Figure A3-6.3: Rotation angle in a short beam

F

Rotation angle

152

The recommended boundary conditions of the thermal model consist of the following:

- The initial temperature of the plate has to be set considering the room temperature of the workshop.

- Convection is recommended to be applied on every area with a film coefficient of 10 W/m2K and a bulk temperature of 25 ºC. Additionally radiation should be considered.

Step 4: The heat input for the thermal model has to be adjusted, preferably by calibrating the real flame used in the process (go to Appendix 2, chapter 5.3). In order to simulate the movement of the flame, it is recommended to impose an offset on the coordinate system origin every load step. The time increment between load steps in the mechanical simulation is recommended to be short enough to allow the heat input applied by the flame to be consistent and continuous enough.

Step 5: The final output from the mechanical simulation basically consists of a bending angle (Figure A3-6.3). If three different short beams with three different applied bending moments have been simulated, a curve that represents the bending angle versus the applied moment can be easily fitted.

Step 6: Finally, by integration of single rotations of the short beams (Figure A3-6.4), the total deflection of the profile and the final shape of the structure can be estimated by using conventional trigonometric relationships.

Figure A3-6.4: Final deflection sketch

153

European Commission

EUR 25120 — Optimisation and improvement of the flame straightening process (Optistraight)

D. Schäfer, V. Rinaldi, D. Beg, P. Može, R. Lacalle, J. Portilla, D. Ferreño, J. A. Álvarez, R. Willms, J. Schütz

Luxembourg: Publications Office of the European Union

2012 — 153 pp. — 21 × 29.7 cm

Research Fund for Coal and Steel series

ISBN 978-92-79-22426-3

doi:10.2777/37733

ISSN 1831-9424

Interested in European research?

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EUROPEAN COMMISSIONDirectorate-General for Research and InnovationResearch Fund for Coal and Steel Unit

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KI-N

A-25120-E

N-N

Flame straightening is a manufacturing process for achieving geometry of elements and members made up of structural steel. The technical background on flame straightening in the workshops of steel constructors is often based only on empirical knowledge, if it exists at all. As a consequence, the straightening of steel construction elements to achieve the required geometrical shape absorbs a large part of the manufacturing costs. This is due to uncertainty about the correct flame straightening process and the lack of background knowledge on its effects. Even though in scientific circles there is some knowledge on the mechanisms of flame straightening for different temperatures, holding times and steel grades, the knowledge is scattered, not well documented and has not been transferred to complex steel structures, sections, stiffness and real (extended) geometries. Also, application techniques, flame straightening procedures and an insight parameter clarification in particular for high strength steels do not exist at all.

This report presents the results of the European research project Optistraight. Through experimental, numerical and analytical investigations, the mechanisms of different flame straightening processes have been clarified. Together with the available, but scattered, knowledge on this fabrication process the results give an in-depth view of the flame straightening process. Based upon this knowledge, prediction means have been developed to estimate the straightening result beforehand and to avoid expensive ‘trial-and-error’ tests, detrimental impacts on the material or excessive energy inputs.

Optimisation and improvement of the flame straightening process

(Optistraight)

doi:2777/37733

Op

timisatio

n and im

pro

vement o

f the flame straig

htening p

rocess (O

ptistraig

ht)E

UE

UR

25120