Search

Welded hollow sections are typical in different industries. A database of fatigue life data of welded hollow section joints covering sequence effects and the accuracy of the linear damage accumulation is presented. The effects of the shape of the applied load spectra, sequence effects of different amplitudes have been investigated using two high-strength steels. This document covers thin-walled tubes of 2-mm thickness made of low-carbon or mild steel 1.8849 (S460MH) and austenitic TWIP-steel 1.4678 + CP700 (X30MnCrN16-14). Constant amplitude and two-level load spectra are presented to check the linear damage accumulation. Using stress concentration factors from finite element analysis, typical FAT classes for the structural and the effective notch stress concepts are checked as well. Both structures show much higher strength compared IIW recommendations by structural stress approach and DVS 0905 by effective notch stress approach. Typical maximum linear damage sums taken from recommendations and codes of 0.2 or 0.5 are exceeded for all spectra investigated and in some of the cases even significantly above 1.0. Transferability of the recommendations to component type structures like those tubular joints made of high-strength steel needs revision to lift its lightweight potential but this will require additional data.

The effective notch stress approach to estimate fatigue strength of welded components requires stress concentration factors of an idealized weld geometry using fictitious notch radii. This paper covers the estimation of those stress concentration factors for welded T-joints in different configurations. In comparison with different existing estimations, new methods using metamodeling by (a) the response surface method based on polynomial regression using coupling terms and (b) based on artificial neural networks are presented. The methods were trained by a large database of reference results obtained by finite element analysis. Besides higher quality of prognosis, the new metamodels enhance the range of allowable parameters of the T-joints compared to existing ones. The resulting methods provide means of obtaining stress concentration factors fast and of sufficient quality which could also be embedded in more complex applications as programmed solutions.

This paper presents the influences of the irregularity “insufficient penetration welding” on the fatigue strength based on experimental and numerical investigations. For this purpose, the fatigue strength of partial penetration butt welds of mild steel (S355) with different non-welded gaps were subjected to fatigue loading. Three different planned gap dimensions (h = 4, 6, 8 mm) at one-sided (Y-seam) and two-sided (DY-seam) partial penetration butt welds were analysed. The results were obtained on a small-scale test specimen with a steel plate thickness of t = 20 mm, without axial and angular misalignment. For the numerical effective notch stress approach, regression formula for efficient analysis of stress concentration factors was determined. The results show that butt joints which are not fully penetrated on one or both sides, and which were previously not permissible for welded joints subject to fatigue loading, can be evaluated in the future on the basis of these investigations.

Stress concentration factors (SCFs) at weld toes and weld roots as required for the effective notch stress concept (see [1, 2]) are usually computed using finite element analysis (FEA) which requires a certain amount of effort for model generation, the solving process, and postprocessing. Regression functions of many FEAs within given parameter bounds provide the possibility of a fast prediction of SCFs. This paper provides new and accurate regression formulae for the estimation of notch stresses at idealized weld geometries on the basis of multiple linear-elastic FEAs for the transverse stiffener (non-load carrying T-joint) under tension and bending of the load carrying slab. Regression of sampled finite element results has been performed using (a) second-order polynomial regression with coupling terms (PRC) and (b) artificial neural networks (ANN). The presented formulae are compared with several existing estimations of stress concentration factors. The new methods appear to show a higher quality of prognosis as well as apply to significant larger ranges of the geometrical parameters of the weld joint. The formulae presented here for the transverse stiffener add another welded joint to a series of similar surrogate models presented from Munich University of Applied Sciences in earlier publications and made available for use by the web-based tool SCF-Predictor.

Polygon meshes and particularly triangulated meshes can be used to describe the shape of different types of geometry such as bicycles, bridges, or runways. In engineering, such polygon meshes can be supplied as finite element meshes, resulting from topology optimization or from laser scanning. Especially from topology optimization, frame-like polygon meshes with slender parts are typical and often have to be converted into a CAD (Computer-Aided Design) format, e.g., for further geometrical detailing or performing additional shape optimization. Especially for such frame-like geometries, CAD designs are constructed as beams with cross-sections and beam-lines, whereby the cross-section is extruded along the beam-lines or beam skeleton. One major task in the recognition of beams is the classification of the cross-section type such as I, U, or T, which is addressed in this article. Therefore, a dataset consisting of different cross-sections represented as binary images is created. Noisy dilatation, the distance transformation, and main axis rotation are applied to these images to increase the robustness and reduce the necessary amount of samples. The resulting images are applied to a convolutional neuronal network.

In-situ manufacturing of thermoplastic composites using the automated fiber placement (AFP) process consists of heating, consolidation and solidification steps. During the heating step using hot gas torch (HGT) as a moving heat source, the incoming tape and the substrate are heated up to a temperature above the melting point of the thermoplastic matrix. The convective heat transfer occurs between the hot gas flow and the composites in which the convective heat transfer coefficient h plays an important role in the heat transfer mechanism which in turn significantly affects temperature distribution along the length, width and through the thickness of the deposited layers. Temperature is the most important process parameter in AFP in-situ consolidation that affects bonding quality, crystallization and consolidation. Although it is well known the convective heat transfer coefficient h is not constant and has a distribution, most studies have assumed a constant value for h for heat transfer analysis which leads to discrepancy between numerical and experimental results. In this study a new function is proposed to approximate the distribution of the convective heat transfer coefficient h in the vicinity of the nip point. Using the proposed convective heat transfer coefficient distribution, a three-dimensional finite element transient heat transfer analysis is performed to predict temperature distribution in the composite parts. An optimization loop is employed to find the free parameters of the distribution function so that the predicted temperature match experimental data. It is shown that, unlike other studies assuming constant h value, not only maximum temperature can be well predicted, but also predicted heating and cooling curves agree well with experimental results. The cooling rate is of significant importance in crystallization behavior and residual stress calculation.

Fatigue evaluation based on the effective notch stress approach requires stress concentration factors for idealized notch geometries. In this paper, stress concentration factors are calculated numerically for different idealizations of the weld geometry. The joints covered here are one-sided Y-butt joints and two-sided DY-butt joints, each with partial and full penetration. The variable parameters of the finite element models are the weld flank angle, the notch radius, the sheet thickness, the ratio of weld seam width to sheet thickness, and the ratio of height of non-fused root face to sheet thickness. Existing estimation formulae for stress concentration factors will be compared with new methods for stress concentration factor estimation: (a) polynomial regression using mixed terms and (b) artificial neural networks. These two methods show similar or superior quality compared with the existing estimations which is expressed through lower estimation errors with respect to the numerically calculated stress concentration factors. The database for the regression analysis using methods (a) and (b) consists of a total of 11,871 design alternatives obtained by finite element analysis. In addition to the improved quality of prognosis, the range of allowable geometrical parameters is significantly increased compared with the existing formulae. The methods (a) and (b) provide a possibility to a fast estimation of stress concentration factors of sufficient quality for a large range of geometrical weld seam parameters. The presented formulae could also be part of a programmed solution.

The automated reverse engineering of wireframes is a common task in topology optimization, fast concept design, bionic and point cloud reconstruction. This article deals with the usage of skeleton-based reconstruction of sketches in 2D images. The result leads to a flexible at least C₁ continuous shape description.

Stress concentration factors (SCF) related to specific geometries, loadings and stress quantities can be calculated efficiently using approximate equations. Those relations are typically based on metamodeling via the results of linear elastic analyses. Existing equations for the, technically speaking, very important circular shafts with shoulder fillets or U-shaped notches under bending, torsion or tension/compression can be appropriated, for example, from the DIN 743-2 : 2012-12. Using these relations requires compliance with underlying constraints and assumptions. For example, equations from that code do not provide SCFs larger than 6 and refer only to the maximum principle stress. Furthermore, DIN 743-2 : 2012-12 does not provide SCFs for combined loading, i. e. simultaneous bending, torsion and tensile/compressive loading. This paper presents new and improved equations based on sampling computational finite element simulations and subsequent regression analyses. Multiple improvements could be achieved: The six equations for circular shafts with shoulder fillets or U-shaped grooves derived from DIN 743-2 : 2012-12 were simplified down to two equations. Despite this simplification, the new equations have been generalized for combined loading and include SCFs larger than 6 as well. Choosing between various equations for each type of loading is no longer necessary. Only one equation for each of the two basic geometries is needed. Moreover, two further equations which refer to von Mises equivalent stress have been added. Although the new equations are easier to handle and allow for an estimation of stress concentration factors larger than 6, the accuracy of the new equations is higher than the accuracy of DIN 743-2 : 2012–12 when compared to SCFs derived from finite element simulation as reference values.

This work includes the application and evaluation of new methods to describe laminate stiffness and strength. Tsai et al. have shown that the trace of laminate stiffness is a rotationally invariant quantity and accurately describes the stiffness potential of a material [1]. In combination with a rotationally invariant strength criterion, the Unit Circle criterion, this allows a simple approach for the dimensioning of fiber composite structures [2]. The use of two biaxial, so called “double-double” [±ϕ/±ψ] sublaminates with the possibility of asymmetrical stacking sequences but homogenization of laminates further simplifies the design and manufacturing process of such structures [3].

Polygon meshes and particularly triangulated meshes can be used to describe the shape of different types of geometry such as bicycles, bridges, or runways. In engineering, such polygon meshes can be supplied as finite element meshes, resulting from topology optimization or from laser scanning. Especially from topology optimization with low member size settings, frame-like polygon meshes with slender parts are typical and often have to be converted into a CAD (Computer-Aided Design) format, e.g., for further geometrical detailing or performing additional shape optimization. Especially for such frame-like geometries, CAD designs are constructed as beams with cross-sections and beam-lines, whereby the cross-section is extruded along the beam-lines or beam skeleton. In our research, automatic parameterization of polygon meshes into a subdivision surface representation is tried out. For this purpose, the beam-lines are approximated by computation of curved skeletons, which are determined by a homotopic thinning method. These skeleton lines are transformed into a subdivision surface control grid by using the Euclidean distance transformation.

This paper covers new and paradigm changing findings in analysis, optimization, design and manufacturing of endless fiber composites as developed by an international group of experts in research and industry inspired and driven by Prof. em. Stephen Tsai. This was presented at the 2019 Munich Lightweight Symposium and is added to the 2020 proceedings as an introduction of the application as presented by Neuhausler in this year’s symposium [7]. The fundamental finding used for the new concepts is the normalizing property of the trace of the stiffness matrix of a unidirectional ply as an invariant material property of plies and laminates. This as a basis enables the use of variable ply angles, a so-called Double-Double architecture of sub-laminates and many more. Much simpler stackings, less design rules to follow, increased strength of laminates, thinner minimum gages, simplified optimization and easier and cost-efficient manufacturing can be obtained.

For the past six years, we have been benefiting from the discovery by Tsai and Melo (2014) that the trace of the plane stress stiffness matrix ( tr ( Q ) ) of an orthotropic composite is a fundamental and powerful scaling property of laminated composite materials. Algebraically, tr ( Q ) turns out to be a measure of the summation of the moduli of the material. It is, therefore, a material property. Additionally, since tr ( Q ) is an invariant of the stiffness tensor Q , independently of the coordinate system, the number of layers, layup sequence and loading condition (in-plane or flexural) in a laminate, if the material system remains the same, tr ( Q ) = tr ( A ∗ ) = tr ( D ∗ ) is still the same. Therefore, tr ( Q ) is the total stiffness that one can work with making it one of the most powerful and fundamental concepts discovered in the theory of composites recently. By reducing the number of variables, this concept shall simplify the design, analysis and optimization of composite laminates, thus enabling lighter, stronger and better parts. The reduced number of variables shall result in reducing the number and type of tests required for characterization of composite laminates, thus reducing bureaucratic certification burden. These effects shall enable a new era in the progress of composites in the future. For the above-mentioned reasons, it is proposed here to call this fundamental property, tr ( Q ) , as Tsai’s Modulus.

Multi-Objective Topology Optimization is a tool for finding low-weight solutions using a discretized geometry of several objective functions. In this contribution, the coupling of heat conduction and elastostatics is covered using the global criteria method. For the comparison of the different objectives and objective distributions typically the single target optimized solutions are selected as normalization criteria [1], [2]. The use of these optimized results requires additional calculation effort, so that this work uses the objective distribution of full material properties instead. For the normalization, the objective distributions will be standardized, and quantile normalized. The normalization strategies are compared to each other by the number of iterations and the resulting objective value.

Commonly, to evaluate the influence of the local weld geometry in fatigue test, small-scale specimens are used, assuming those represent a longer weld adequately. In this study, a comparison between short specimens and a long weld is performed. A method is developed for the statistical evaluation of weld toe radii and angles, stress concentration factors and weld quality classes. The results show a strong sampling rate dependence and lower ISO 5817:2014 weld quality results for higher sampling rates. Comparable results between short specimens and a long weld can be achieved using modal values of the parameters assuming a lognormal distribution.

Feature lines such as sharp edges are the main characteristic lines of a surface. These lines are suitable as a basis for surface reconstruction and reverse engineering [1]. A supervised deep learning approach based on graph convolutional networks on estimating local feature lines will be introduced in the following. We test this deep learning architecture on two provided data sets of which one covers sharp feature lines and the other arbitrary feature lines based on unnoisy meshed constructive solid geometry [CSG]. Furthermore. we use a data balancing strategy by classifying different feature line types. We then compare the selected architecture with classical machine learning models. Finally. we show the detection of these lines on noisy and deformed meshes.

The effective notch stress approach for fatigue assessment of welded joints according to the IIW-recommendations is a widely accepted procedure and has already become part of many technical rules and guidelines. The method provides additional means for evaluation of welded joints besides classical nominal and structural stress concepts. In contrast to the latter concepts, effects of many local weld parameters can only be studied analytically using the effective notch stress approach. The effective notch stress approach requires use of a standardized idealization of the weld detail and use of continuum mechanics theory for stress analysis (i.e. use of solid elements for finite element models). The idealization with respect to geometry includes modeling of all weld toes and weld roots using a reference radius together with an associated S-N curve. For joints facing low stress concentration, like butt welds or welds at thin sheets, the method requires checking the fatigue behavior of the base material as well. This sometimes is not considered in using the effective notch stress method, because it is not so obvious. Explanation of background information and a reanalysis of an example from literature is presented here for explanation and demonstration of this concept and of specific needs for welds with low stress concentration. Also the selection of the size of the reference radius will be discussed in this paper by using different radii for the case investigated here. Finally the application limits of the S-N curves as given in the IIW-recommendations is discussed since the selected example is slightly out of the range of application of the S-N curves.

The main aim of the project was to evaluate the influence of combined effects of fatigue loading and exposure to cyclic corrosion testing on the corrosion and the fatigue resistances of coated thin sheet steel joined by spot welding. Seven types of steel including cold rolled mild steels, high-strength steels, and press-hardened steels (PHS) were selected and provided by steel suppliers for fatigue strength evaluation of resistance spot welding (RSW) assemblies. Panels were joined using conventional resistance spot welding in both lap-shear and T-peel designs. Joined samples were painted by e-coating following the industrial process. Metallographic characterization of the steels revealed that microstructures and metallic coating composition and thickness were as expected. Cross-section of spot welds showed good quality and typical evolutions of hardness. Tensile tests performed on reference samples (non-exposed to corrosion) showed tensile which were obviously configuration and material dependent. For each configuration, three replicates were tested and did not show significant scatter. The results were and will be used to evaluate the influence of corrosion on tensile strength of the joined samples. Fatigue tests were performed “in air”, meaning without corrosion exposition, leading to typical SN-lines. The results will be used to evaluate the influence of corrosion on the fatigue resistance of the joined samples.

The main aim of the project was to evaluate the influence of combined effects of fatigue loading and exposure to cyclic corrosion testing on the corrosion and the fatigue resistances of coated thin ...

The aim of this study is to evaluate fatigue performance of joined assemblies (spot weld and/or adhesive bonding) in corrosive environment. Various assemblies have been tested in alternated and simultaneous fatigue-corrosion modes. Adhesive joints are strongly affected by simultaneous fatigue-corrosion with a large drop of the fatigue life compared to results in air. By alternating fatigue and corrosion, the reduction of fatigue life is important. For spot welding, fatigue life is decreased at higher load amplitudes and increased at lower amplitudes. These results are strongly linked to the opening of the gap near the spot weld at high load amplitudes. At low amplitudes, corrosion might limit the local stress at the notch root of the weld.

As elastic–plastic fatigue analyses are still time consuming the simplified elastic–plastic analysis (e.g. ASME Section III, NB 3228.5, the French RCC-M code, paragraphs B 3234.3, B 3234.5 and B3234.6 and the German KTA rule 3201.2, paragraph 7.8.4) is often applied. Besides linearly elastic analyses and factorial plasticity correction (Ke factors) direct methods are an option. In fact, calculation effort and accuracy of results are growing in the following graded scheme: a) linearly elastic analysis along with Ke correction, b) direct methods for the determination of stabilized elastic–plastic strain ranges and c) incremental elastic–plastic methods for the determination of stabilized elastic–plastic strain ranges. The paper concentrates on option b) by substantiating the practical applicability of the simplified theory of plastic zones STPZ (based on Zarka's method) and – for comparison – the established Twice-Yield method. The Twice-Yield method is explicitly addressed in ASME Code, Section VIII, Div. 2. Application relevant aspects are particularly addressed. Furthermore, the applicability of the STPZ for arbitrary load time histories in connection with an appropriate cycle counting method is discussed. Note, that the STPZ is applicable both for the determination of (fatigue relevant) elastic–plastic strain ranges and (ratcheting relevant) locally accumulated strains. This paper concentrates on the performance of the method in terms of the determination of elastic–plastic strain ranges and fatigue usage factors. The additional performance in terms of locally accumulated strains and ratcheting will be discussed in a future publication.

Organisationen und Prozesse in der Fahrzeugentwicklung sind in den letzten Jahrzehnten vordringlich auf standardisierte Ablaufe einer Anderungskonstruktion ausgerichtet wurden. Simulationsunterstutze Entwicklung wurde vor allem im Hinblick auf die Steigerung der Prognosegute und standardisierter Simulationsprozesse optimiert. Eine Analyse des Entwicklungsprozess in der fruhen Phase bei einem grosen deutschen OEM (Damjanovic et al. 2014) hat ergeben, dass in Verbesserungen des Wissens- und Datenmanagements sowie der Modellierung von (nicht nur) konzeptioneller Geometrie inkl. Simulation auf 3D sowie die Verbesserung interdisziplinarer Zusammenarbeit erhebliche Potenziale schlummern. Losungsansatze ergeben sich hierzu beispielsweise im Aufbrechen einer strikten Trennung von fruher und spater Phase, in Modellierungstechniken, die durch Parametrisierung Variantenvielfalt fordern, durch Zugriff und damit Sicherstellung der Nachvollziehbarkeit des sich steigernden Detaillierungsgrads uber die PLM/SPDM-Systeme, in einer verbesserten simulations- und optimierungsgesteuerten Konstruktion mit ganzheitlich greifenden, d.h. holistischen Ansatzen, oder auch durch Verbesserung unternehmensweiter, effizienter Datenubertragung und einer Verminderung von Barrieren bzw. Interfaces.

Durch das zunehmende oekologische Bewusstsein der Bevoelkerung und die Vorgaben der Politik und Gesetzgebung steigt der Kundenwert von Gewichtseinsparungen am Fahrzeug. Es wird aufgezeigt, dass eine Gewichtsreduzierung zugunsten der kundenwerten Eigenschaften haeufig mit hoeheren Leichtbaukosten verbunden ist. Der Beitrag beschreibt in diesem Zusammenhang die negative Gewichtsspirale. Darunter versteht man den Kreislauf, bei dem durch eine Funktionserweiterung die Anforderungen an ein einzelnes Bauteil im Fahrzeug steigen und damit als Primaereffekt oft auch dessen Gewicht zunimmt. Um dieses schwerere Bauteil bei gleichen Gesamtfahrzeugeigenschaften zu integrieren, muss die Peripherie ertuechtigt werden, was zu weiteren Gewichtsmehrungen, den so genannten Sekundaereffekten, fuehrt. Es wird diesbezueglich ein holistischer Ansatz vorgestellt, der - um im Gewichtsprozess einer Fahrzeugentwicklung den Aufwand fuer primaere Leichtbaumassnahmen zu reduzieren - Sekundaereffekte moeglichst vollumfaenglich ausschoepft. Der Ansatz macht sich damit die Gewichtsspirale als gezielte Massnahme zum kostenguenstigen Leichtbau zunutze. Die jeweiligen Prozessschritte des Ansatzes - naemlich die Festlegung des strategischen Zielgewichts, die Vorwegnahme der Sekundaereffekte (durch Korrelationsanalysen), die Ableitung erforderlicher Primaermassnahmen, die Festlegung der Auslegungsgewichte sowie die konkrete Auslegung des Fahrzeugs - werden anhand eines Rechenbeispiels vorgestellt. Der Beitrag schliesst mit dem Ausblick, dass die "negative Gewichtsspirale" von Vorteil sein kann, sofern es gelingt, den Gewichtsprozess in der Fahrzeugentwicklung stringent darauf auszurichten.

In the context of an internal study, performed by the BMW Group, dealing with the potential of secondary mass effects, several differentiated correlation analysis have been completed in order to determine the total vehicle mass sensitivity of several parts and component assemblies. The method applied in this study is similar to the approach of the negative weight spiral by Braess [5], but expanded with the secondary mass effects of the car body. For the study, two identical designed cars, the BMW 5-series and 7-series of the current and predecessor generation, were compared, ❷. Correlation coefficients in dependency of the total vehicle weight were determined for the components chassis KF, drive train KA and car body KRK , Eq. 1. In a first approach, the vehicle interior could be treated as non-related to the total vehicle mass and was therefore not relevant for secondary mass savings. This assumption is applied in other publications as well. The drive train sensitivity to the total vehicle weight could be determined by the correlation of the drive train mass to the torque output. Secondary mass effects in the drive train can only be utili­zed if the primary mass effect is not transferred into increased driving performance. Open image in new window ❷ Correlation using the example of chassis mass in relation to the total vehicle mass The component specific correlation coefficients KF = 0.15 , KRK = 0.16 and KA = 0.06 add up to a total vehicle correlation coefficient for secondary mass effects of K = 0.37. As a result, a mass reduction of ∆G = 100 kg yields secondary mass effects of ∆Gsec = 37 kg. In case of need, this correlation analysis could be refined to the effect of single parts. Therefore, a procedure according to the pareto principle makes sense since 20 % of the vehicle parts include approximately 80 % of the total vehicle mass. The determined secondary effects again initiate new secondary effects and trigger a chain reaction. As a limit value it can be shown that a primary impulse of ∆G = 100 kg combined with the correlation co­­efficient K = 0.37 leads to a potential in total mass reduction of ∆Gtotal = 159 kg, Eq. 2. This maximum secondary effect then corresponds to 37 % of the total mass reduction and requires a primary impulse of 63 % of the total mass savings. With respect to the development pro­cess of a car, the feasible secondary mass potential essentially depends on the point in time a primary impulse is activated. To achieve the full secondary mass potential, the dimensioning framework has to be specified before the design of any part will be done or any supplier query will be requested. Depending on the manufacturer, the time window of this early concept phase ends approximately 50 to 55 months before start of production. A detailed picture of the time based dependency of possible secondary mass effects could be obtained if each part’s construction completion is weighted with its corresponding part mass along the development timeline, ❸. It is noticeable in the mentioned figure that secondary effects barely could be activated from 45 months before start of production. Open image in new window ❸ Potential secondary mass effects and their dependency on the implementing point in the product ­development process A holistic approach was developed considering the aforementioned statements and to minimize the efforts for primary lightweight actions by implementing secondary mass effects as numerous as possible. This process in particular is efficient in evolutionary car development (i.e. if a successor of an established vehicle is developed). The proceeding involves the following steps of the process: : determining a strategic target weight : anticipation of secondary mass effects : derivation of the required primary impulse : definition of the required mass of all parts/subsystems : dimensioning and design of the vehicle. $$ {\rm{K}}_{\rm{F}} = \frac{{{\rm{Delta}}\,{\rm{chassis}}\,{\rm{mass}}}}{{{\rm{Delta}}\,{\rm{total}}\,{\rm{vehicle}}\,{\rm{mass}}}} = \frac{{\Delta {\rm{G}}_F }}{{\Delta {\rm{G}}}} = 0.15 $$ (Eq. 1.1) $$ {\rm{K}}_{{\rm{RK}}} = \frac{{{\rm{Delta}}\,{\rm{body}}\,{\rm{structure}}\,{\rm{(without}}\,{\rm{closures)}}}}{{{\rm{Delta}}\,{\rm{total}}\,{\rm{vehicle}}\,{\rm{mass}}}} = \frac{{\Delta {\rm{G}}_{{\rm{RK}}} }}{{\Delta {\rm{G}}}} = 0.16 $$ (Eq. 1.2) $$ \begin{array}{rll} {\rm{K}}_{{\rm{RK}}} &=& \frac{{{\rm{Delta}}\,{\rm{drive}}\,{\rm{train}}\,{\rm{mass}}}}{{{\rm{Delta}}\,{\rm{torque}}\,{\rm{output}}}} \cdot \frac{{{\rm{Proportionaly}}\,{\rm{factor}}\,{\rm{according}}\,{\rm{to}}\,{\rm{W}}{\rm{.}}\,{\rm{Stork}}\,{\rm{[3]}}}}{{{\rm{Average}}\,{\rm{acceleration}}\,{\rm{time}}\,{\rm{0}}\,{\rm{to}}\,{\rm{100}}\,{\rm{km/h}}}}\\ &=& \frac{{\Delta {\rm{G}}_{\rm{A}} }}{{\Delta {\rm{G}}}} = \frac{{1.35\frac{{\rm{s}}}{{{\rm{m}}^{\rm{2}} }}}}{{{\rm{t}}_{{\rm{0 - 100}}\,{\rm{k/mh}}} }} = 0.06 \end{array}$$ (Eq. 1.3) $$ \begin{array}{rll} \rm{Total \, secondary \, mass \, potential \, after \, n \, iterations:} \\&&\Delta {\rm{G}}_{{\rm{ges}}} = \Delta {\rm{G}} + \Delta {\rm{G}} \cdot {\rm{K}} + \Delta {\rm{G}} \cdot {\rm{K}}^2 + \ldots + \Delta {\rm{G}} \cdot {\rm{K}}^{\rm{n}} = \mathop \sum \limits_{{\rm{i}} = {\rm{l}}}^{\rm{n}} \Delta {\rm{G}} \cdot {\rm{K}}^{\rm{i}} = \\ && \frac{{\Delta {\rm{G}}({\rm{Kn}} - {\rm{1}})}}{{{\rm{K}} - 1}};{\rm{with}}\,{\rm{n}} \to \infty \,{\rm{and}}\,\Delta {\rm{G}} = 100\;{\rm{kg}}\,{\rm{and}}\,{\rm{k}} = 0.37 \\&&{\rm{result}}\,{\rm{in}}\,{\rm{a}}\,{\rm{limit}}\,{\rm{value}}\,{\rm{of:}}\lim _{{\rm{n}} \to \infty } \left( {\frac{{\Delta {\rm{G}}\left( {{\rm{K}}^{\rm{n}} - 1} \right)}}{{{\rm{K}} - 1}}} \right) = 159\;{\rm{kg}} \end{array} $$ (Eq. 2) $$ \begin{array}{rll} &&{\rm{x}} \cdot {\rm{costs}}\,{\rm{of}}\,{\rm{conventional}}\,{\rm{lightweight}}\,{\rm{design}} = {\rm{costs}}\,{\rm{of}}\,{\rm{lightweight}}\,{\rm{design}}\,{\rm{with}} \\&&{\rm{secondary}}\,{\rm{mass}}\,{\rm{effects}}\,{\rm{x}} \cdot \left[ {\Delta {\rm{G}} \cdot {\rm{5}}\frac{{{€}}}{{{\rm{kg}}}}} \right] = \left( { - 0.37 \cdot \Delta {\rm{G}}} \right) \cdot 1.4\frac{{{€}}}{{{\rm{kg}}}} + \left( {\Delta {\rm{G}} - 0.37 \cdot \Delta {\rm{G}}} \right) \cdot 5\frac{{{€}}}{{{\rm{kg}}}}\\&&{\rm{x}} = 53\% \end{array} $$ (Eq. 3) The determination of the strategic target weight should be based on costumer related characteristics, requirements and on business objectives affected by the total vehicle mass. To support the feasibility of secondary mass effects, it is essential to imply the existing discontinuous weight functions systematically in the dimensioning processes. Moreover, the timeline is critical for the success of this process. It is important to define and communicate the dimensioning masses derived from the strategic target weight. This has to be done before any major detailed design is finished in order to utilize the total potential of secondary mass effects including all iterations possible. For the next step, all possible secondary mass effects, as obtained by applying the component specific correlation coefficients and including all possible iterations, have to be subtracted from the difference of the initial weight to the target weight. The total potential of all possible iterations can only be achieved if the chain of effects is fully understood and all partners involved adopt the proper attitude for this process. The remaining difference between the weight loss already achieved and the target value defines the essential and critical demand of primary mass effects in form of lightweight measures. For a weight re­duction of 100 kg in total, 37 kg of secondary mass effects can be achieved if all iterations are included. The remaining 63 kg have to be achieved by activating primary mass effects, ❹. All requirements, the total vehicle target weight, the component allocated secondary mass effects, the needed primary mass impulse and the resulting dimensioning masses on total vehicle, system and component level have to be agreed on with all process members before launch of detailed designs. Based on these premises each component has to be dimensioned and designed. Open image in new window ❹ Course of action to implement a holistic approach of weight reduction In order to evaluate the economic potential of this approach a comparison with a conventional material lightweight strategy can be done. For the conventional lightweight approach, cost savings of approximately 5 € per kg are assumed [6]. With the discussed utilization of secondary mass effects some material (0.37∙∆G) can be saved. In doing so, this mass saving could be capitalized with assumed, but absolute­­ly realistic average material costs of 1.40 €/kg. Only the lightweight costs for the re­­maining primary impulse (∆G – 0.37∙∆G) have “to be paid for”. In total, this holistic approach implies the potential to cut the lightweight costs of conventional lightweight design in half, Eq. 3. Summing up, the negative weight spiral could definitely be a real blessing if the process for vehicle development with focus on lightweight design is rearranged successfully. Using an early determination of a strategic target weight, anticipation of secondary mass effects and an agreement of required subsequent mass savings, the needed primary lightweight effort of the targeted weight reduction could be re­­duced by 37 %. As a result of this ap­­proach, the cost for lightweight expenses in manufacturing could be reduced by almost 50 %.

Analysis of welded structures still remains a challenge for the analyst and in fact cannot be considered as fully solved for practical applications. For many years, a large international aggregation of researchers has developed methods to assess fatigue behaviour of welded structures. Nowadays many suggestions and methods exist to estimate fatigue life of welded structures with respect to nominal, structural, notch stress or fracture mechanics approaches. All of them are still under improvement. The high motivation and many activities of experts in the International Institute of Welding (IIW) group of researchers is a good demonstration of the complexity and need for analysis methods in that field. The purpose of this paper is to provide some discussion on selected methods available. Both authors are giving lectures to transfer methods to industrial applications. It is their experience that a large amount of knowledge has been developed although proper applications require some grading and comments on the use of those methods. This paper should give some comments and recommendations for the practical application of a selection of methods already available. A hierarchical two-step procedure for the assessment of large welded structures will be described and recommended. Also benchmark results are presented on a sample structure for sake of comparison of a few selected methods. Finally a presentation of results obtained by application of selected methods on real structures in comparison with fatigue lives from experiments will be presented. The methods selected within the paper cover the approaches for modelling, structural analysis and assessment of welded structures using finite element analysis (FEA) and stress based concepts for fatigue life estimation.

Kurzfassung Die Analyse geschweißter Strukturen ist eine Herausforderung für den Berechnungsingenieur. Verbesserte Soft- und Hardwareleistung ermöglicht heute die Berechnung ganzer Baugruppen in einer Analyse. Gerade für verschweißte Strukturen bieten sich durch diese Möglichkeiten besondere Vorteile, die zu einer genaueren Abbildung der Kraftflüsse in der Struktur beitragen und zudem dem Berechnungsingenieur die Arbeit erleichtern. Der vorliegende Beitrag behandelt die Vorgehensweise zur Modellbildung, Strukturanalyse und Auswertung von geschweißten Strukturen mit der Finite-Elemente-Methode. Es wird ein hierarchisches Konzept bestehend aus einer Globalanalyse eines Schweißteils zur Lokalisierung kritischer Positionen anhand von Strukturspannungen und lokalen Betrachtungen nach Kerbspannungskonzept beschrieben.

In der Automobilindustrie werden Karosserien inzwi- schen auch ohne physikalische Prototypen hinsichtlich der Kriterien der Lackierung und des Korrosionsschutzes umfassend virtuell gepruft. Mit neuen Software-Paketen ist die Simulation der thermischen Trocknung, der UV-Hartung, der KTL-Beschichtung, der Hohlraum- versiegelung sowie von Tauchprozessen moglich.

As elastic-plastic fatigue analyses are still time consuming the simplified elastic-plastic analysis (e.g. ASME Section III, NB 3228.5, the French RCC-M code, paragraphs B 3234.3, B 3234.5 and B3234.6 and the German KTA rule 3201.2, paragraph 7.8.4) is often applied. Besides linearly elastic analyses and factorial plasticity correction (Ke-factors) direct methods are an option. In fact, calculation effort and accuracy of results are growing in the following graded scheme: a) linearly elastic analysis along with Ke correction, b) direct methods for the determination of stabilized elastic-plastic strain ranges and c) incremental elastic-plastic methods for the determination of stabilized elastic-plastic strain ranges. The paper concentrates on option b) by substantiating the practical applicability of the simplified theory of plastic zones STPZ (based on Zarka’s method). Application relevant aspects are particularly addressed. Furthermore, the applicability of the STPZ for arbitrary load time histories in connection with an appropriate cycle counting method is discussed. Note, that the STPZ is applicable both for the determination of (fatigue relevant) elastic-plastic strain ranges and (ratcheting relevant) locally accumulated strains. This paper concentrates on the performance of the method in terms of the determination of elastic-plastic strain ranges and fatigue usage factors. The additional performance in terms of locally accumulated strains and ratcheting will be discussed in a future publication.