Flanging pre-compression and thermal management bushing cae segmented stiffness design method for a mating structure

Abstract

The present application relates to the technical field of simulation design of automobile rubber damping parts, in particular to a CAE segmented stiffness design method for a heat management bushing with a flange pre-compression and a matched structure, a finite element model of a two-half matched split bushing is established, containing a metal inner and outer tube, a main body rubber and a flange; an axial compression load is applied to simulate the flange pre-compression, and a pre-compression stress field is obtained as an initial condition for stiffness calculation; the radial stiffness is solved first to lock the main body parameters, then the axial stiffness is extracted by segmenting and extracting the axial stiffness with displacement threshold as the boundary under continuous axial loading; the flange thickness is taken as a variable for iterative optimization, the optimal thickness is selected and the compression amount is calculated reversely; and the shaped parameters are output. The method realizes segmented stiffness control, radial-axial decoupling, improves accuracy and shortens the cycle.

Description

Technical Field

[0001] This invention relates to the field of simulation design technology for automotive rubber vibration damping components, and in particular to a CAE segmented stiffness design method for thermal management bushings of flanged pre-compression and mating structures. Background Technology

[0002] Thermal management bushings are key vibration isolation and support components of the thermal management module in new energy vehicles. Their axial stiffness characteristics directly affect the overall vehicle's NVH performance and the module's installation stability. Currently, these bushings mostly adopt a one-piece vulcanized molding structure. After designers determine initial parameters such as rubber hardness and wall thickness based on experience, they obtain stiffness curves through mold making, prototyping, and bench testing. If the requirements are not met, the structure is repeatedly modified and re-produced until the requirements are met. This development process is highly dependent on physical testing, usually requiring multiple rounds of mold making and prototyping, resulting in a long cycle and high cost, which is no longer able to meet the rapid iteration development needs of new energy vehicles.

[0003] With increasingly stringent requirements for bushing performance, a new type of bushing with a pre-compression structure featuring a two-part mating end ring flange has emerged in recent years. This type of structure utilizes two split bushings axially opposed to each other, employing the pre-compression deformation of the end flanges to provide low-stiffness buffering under small displacements, while the main rubber layer provides stable support under large displacements, thus achieving segmented stiffness characteristics. However, existing simulation analysis methods lack a dedicated modeling and simulation process for this special structure. Designers cannot accurately simulate the pre-compression assembly state of the flanges and its contribution to axial stiffness, resulting in significant discrepancies between simulation predictions and actual measurements, thus limiting the guiding value of the simulation process.

[0004] In addition, the radial stiffness and axial stiffness of traditional integrated bushings are highly coupled, making it difficult for designers to adjust the axial segment stiffness separately without affecting the radial support performance. The determination of key parameters such as flange thickness and pre-compression amount still relies on manual trial and error, lacking a standardized parameter iteration optimization process, resulting in low efficiency and strong reliance on experience throughout the development process, making it difficult to form reusable design knowledge.

[0005] Therefore, how to establish a CAE simulation design method for two-half assembled flanged pre-compression thermal management bushings that can realistically simulate the flanged pre-compression assembly state, achieve radial and axial segmented stiffness decoupling optimization, and have a standardized parameter iteration process is a technical problem that urgently needs to be solved in this field. Summary of the Invention

[0006] The purpose of this invention is to provide a CAE segmented stiffness design method for thermal management bushings of flanged pre-compression and mating structures, which realizes precise and controllable segmented stiffness and decoupling of radial and axial performance, significantly improving simulation accuracy and shortening the development cycle.

[0007] To achieve the above objectives, the present invention provides a CAE segmented stiffness design method for thermal management bushings of flanged pre-compression and mating structures, comprising the following steps: A three-dimensional finite element simulation model of a two-part assembled split bushing is established. The split bushing includes an inner metal tube, an outer metal tube, a main rubber layer located between the two, and an annular flange structure set at the axial end of the main rubber layer. An axial assembly clamping load is applied to the three-dimensional finite element simulation model to drive the two split bushings to move towards each other, causing the two annular flange structures to undergo pre-compression deformation. The pre-compression stress field is obtained by solving the problem and is set as the initial stress field. Based on the initial stress field, a radial displacement load is applied to the inner metal tube, the radial stiffness curve is obtained by solving, and the structural parameters of the main rubber layer that meet the radial support requirements are locked accordingly. After locking the structural parameters of the main rubber layer, axial displacement load is continuously applied along the axial direction, and the entire axial loading stroke is divided into the first axial displacement interval and the second axial displacement interval with the preset displacement threshold as the dividing point. The axial displacement-reaction data of each interval are extracted to obtain the segmented axial stiffness curve. Using the original thickness of the annular flange structure as the iterative variable, multiple simulation models with different original flange thicknesses are established. The pre-compression simulation and segmented stiffness solution are repeatedly executed. The segmented axial stiffness curves obtained from each simulation are compared with the preset design target curves to select the optimal original flange thickness. The corresponding assembly pre-compression amount is then calculated in reverse. The finalized structural parameters, including the optimal original flange thickness and assembly pre-compression, are output to the process end.

[0008] When establishing the three-dimensional finite element simulation model, two axially symmetrical and structurally identical split bushing geometric entities are constructed. The two split bushings are arranged in the same global coordinate system in a coaxial alignment manner, and their annular flange end faces each other in the axial direction, forming a two-half assembled overall structural form.

[0009] The inner and outer metal tubes are modeled using linear elastic constitutive models, with their respective elastic moduli and Poisson's ratios input. The main rubber layer and the annular flange structure are modeled using hyperelastic constitutive models, with hyperelastic material constants determined by fitting mechanical test data of standard rubber specimens. The main rubber layer and the annular flange structure are given identical hyperelastic material parameters.

[0010] Specifically, the inner circumferential surface of the main rubber layer and the outer circumferential surface of the metal inner tube, as well as the outer circumferential surface of the main rubber layer and the inner circumferential surface of the metal outer tube, are both set as adhesive binding constraints to simulate the integral fixed connection state of vulcanization bonding; the annular flange end faces that are opposite each other on the two split bushings are set as face-to-face contact pairs, and normal hard contact behavior and tangential Coulomb friction contact behavior are set to simulate the assembly contact and relative sliding state when the two halves of the bushing are assembled.

[0011] In the process of applying an axial assembly clamping load to simulate the pre-compression of the flange, a completely fixed constraint is applied to the outer metal tube of one split bushing, and an axial forced displacement driving load is applied to the outer metal tube of the other split bushing, causing the two split bushings to move towards each other along the axial direction until the thickness of the annular flange structure is compressed from the free state to the target assembly thickness. After the solution is completed, the deformation configuration and stress field distribution data of the flange region under the target compression thickness are extracted, and the stress field is written into the initial stress field for subsequent stiffness calculation.

[0012] The value of the preset displacement threshold is equal to the value of the single-sided pre-compression of the annular flange structure during the assembly process.

[0013] In the axial segmented stiffness simulation, within the first axial displacement range from zero to a preset displacement threshold, the axial reaction force is contributed by the bending and compressive deformation of the annular flange structure. The average stiffness of this range is extracted to match the small-amplitude vibration buffer target. Within the second axial displacement range from the preset displacement threshold to the maximum axial displacement, the two annular flange end faces are out of contact and squeezed, and the axial load is mainly borne by the main rubber layer. The average stiffness of this range is extracted to match the set displacement limit support target.

[0014] Among them, using the original thickness of the annular flange structure as the iterative variable, multiple simulation models with different original flange thicknesses are established, including: Using the original thickness of the annular flange as the sole iterative variable, while keeping all dimensional parameters and material properties of the inner metal tube, outer metal tube, and main rubber layer unchanged, a deviation value is taken on each side of the preset center value to form at least three simulation schemes with different original flange thicknesses. The target compression thickness of the flange after assembly remains consistent in each scheme.

[0015] When comparing and screening the segmented axial stiffness curves obtained from each group of simulations with the preset design target curves, a comprehensive evaluation is conducted based on the following three criteria: whether the error between the average stiffness of the simulation curve and the target curve in the first axial displacement interval falls within the preset allowable range; whether the error between the average stiffness of the simulation curve and the target curve in the second axial displacement interval falls within the preset allowable range; and whether the stiffness transition of the simulation curve at the preset displacement threshold boundary point meets the continuity requirement.

[0016] In this process, after determining the optimal original flange thickness, the pre-compression amount for assembly is calculated by subtracting the target compression thickness after assembly from the optimal original flange thickness to obtain the pre-compression amount on one side. At the same time, the assembly clamping force corresponding to this compression amount is extracted from the pre-compression simulation results. The original flange thickness, the target compression thickness after assembly, the pre-compression amount on one side, and the assembly clamping force are all output as the finalized structural parameters to the process end.

[0017] This invention discloses a CAE segmented stiffness design method for a thermal management bushing with flange pre-compression and assembly structure. A three-dimensional finite element simulation model of the two-part assembled bushing is established. The bushing includes an inner metal tube, an outer metal tube, a main rubber layer, and an end annular flange. An axial assembly clamping load is applied to the model to simulate the flange pre-compression condition and obtain the pre-compression stress field, which is used as the initial stress field for subsequent stiffness calculations. Based on the initial stress field, the radial stiffness is first solved and the main structural parameters are locked. Then, a load is continuously applied along the axial direction, and the axial stiffness curve is extracted segmentally using a preset displacement threshold as the dividing point. Multiple schemes are compared and optimized using the original flange thickness as an iterative variable to select the optimal flange thickness and calculate the assembly pre-compression amount in reverse. The finalized structural parameters are output to the manufacturing process. This invention achieves precise and controllable segmented stiffness and decoupling of radial and axial performance, significantly improving simulation accuracy and shortening the development cycle. Attached Figure Description

[0018] To more clearly illustrate the technical solutions in the embodiments of this application or the prior art, the accompanying drawings used in the description of the embodiments or the prior art will be briefly introduced below.

[0019] Figure 1 This is a schematic diagram illustrating the steps of a CAE segmented stiffness design method for a heat management bushing of a flanged pre-compression and mating structure provided by the present invention.

[0020] Figure 2 This is a flowchart illustrating a CAE segmented stiffness design method for a heat management bushing of a flanged pre-compression and mating structure provided by the present invention. Detailed Implementation

[0021] Exemplary embodiments will now be described in detail, examples of which are illustrated in the accompanying drawings. When the following description relates to the drawings, unless otherwise indicated, the same numbers in different drawings represent the same or similar elements. The embodiments described in the following exemplary embodiments do not represent all embodiments consistent with this application.

[0022] The terminology used in this application is for the purpose of describing particular embodiments only and is not intended to be limiting of the application. The singular forms “a,” “the,” and “the” used in this application and the appended claims are also intended to include the plural forms unless the context clearly indicates otherwise. It should also be understood that the term “and / or” as used herein refers to and includes any or all possible combinations of one or more of the associated listed items.

[0023] It should be understood that although the terms first, second, third, etc., may be used in this application to describe various information, such information should not be limited to these terms. These terms are only used to distinguish information of the same type from one another. For example, without departing from the scope of this application, first information may also be referred to as second information, and similarly, second information may also be referred to as first information. Depending on the context, the word "if" as used herein may be interpreted as "when," "when," or "in response to determination."

[0024] Please see Figures 1-2 This invention provides a CAE segmented stiffness design method for thermal management bushings of flanged pre-compression and mating structures, comprising the following steps: S1. Establish a three-dimensional finite element simulation model of the two-part assembled split bushing. The split bushing includes a metal inner tube, a metal outer tube, a main rubber layer located between the two, and an annular flange structure set at the axial end of the main rubber layer.

[0025] Specifically, firstly, based on the 3D solid data in the product design drawings, two structurally identical and axially symmetrical split bushing geometric entities are independently constructed. Each split bushing adopts a sleeve-shaped rotating structure design, which consists of three core components radially from the inside out: the innermost layer is a cylindrical metal inner tube, used to mate with the mounting mandrel of the thermal management module; the outermost layer is a cylindrical metal outer tube, used to mate with the vehicle's mounting bracket or suspension support; located between the inner and outer metal tubes is the main rubber layer, which radially fills the annular space between the inner and outer metal tubes and extends axially to the end region of the bushing.

[0026] Specifically, at the axial end of the main rubber layer, the rubber material extends continuously and bulges outward in the radial direction, forming an annular flange structure with a specific free thickness. This annular flange protrudes outward in a cantilever shape in the axial section, and its outermost radial end can extend to the vicinity of the outer circumference of the metal outer tube or even slightly beyond the outer tube, in order to provide pre-compression deformation and axial buffering functions during subsequent assembly.

[0027] After completing the independent geometric modeling of the two separate bushings, based on the actual assembly constraints, the two separate bushings are arranged coaxially aligned in the same global coordinate system, so that their axes coincide. At the same time, the two separate bushings are set to face each other axially, with their flanged end faces facing each other. That is, the annular flanged end faces of the first separate bushing and the annular flanged end faces of the second separate bushing axially and are separated by a very small initial gap, thus forming a complete two-half assembled overall structure.

[0028] Once the geometric model is constructed, the corresponding material mechanical properties are assigned to the components of different materials in the model. It is necessary to strictly distinguish between metal components and rubber components and use different constitutive models for description.

[0029] For both inner and outer metal tubes, since the deformation under normal operating loads is minimal and falls within the linear elastic range, their material properties are defined using a linear elastic constitutive model. Specifically, the elastic modulus parameter of the metal material needs to be input to characterize its ability to resist tensile and compressive deformation, while the Poisson's ratio parameter needs to be input to characterize the proportional relationship between lateral and longitudinal deformation. This linear elastic model can effectively reduce the computational burden of subsequent nonlinear solutions while ensuring computational accuracy.

[0030] For the main rubber layer and the annular flange structure with identical material properties, given the significant material and geometric nonlinear characteristics of rubber under large deformation conditions, and the complex nonlinear functional relationship between stress and strain, its material properties are defined as a hyperelastic constitutive model. In practice, the material constants of the hyperelastic constitutive model need to be determined by fitting basic mechanical test data such as uniaxial tension, plane shear, and equal biaxial tension of standard rubber specimens. Common hyperelastic model forms, such as the Mooney-Rivlin model or the Ogden model, can be used, with the corresponding strain energy density function coefficients input to accurately describe the nonlinear mechanical response behavior of rubber under multi-directional complex loads such as axial compression, radial extrusion, and flange bending. It should be particularly noted that although the main rubber layer and the annular flange structure are integrally vulcanized and use the same hyperelastic material parameters, the model definition must ensure that this set of material parameters is simultaneously assigned to both geometric regions.

[0031] After defining the materials, accurate connection and contact constraints need to be established between the various components of the model to simulate the physical state of actual vulcanization bonding and assembly contact. This step is divided into two categories based on the different interaction properties: adhesive bonding contact and face-to-face mechanical contact.

[0032] The first type is the adhesive bonding relationship between rubber and metal. For each separate bushing, the inner circumferential surface of its main rubber layer and the outer circumferential surface of the metal inner tube form a strong chemical bond during the actual vulcanization process, and the two do not experience any relative sliding or separation during service. Similarly, the outer circumferential surface of the main rubber layer and the inner circumferential surface of the metal outer tube are also fixedly connected through vulcanization bonding. To accurately simulate this physical state in the finite element model, adhesive bonding constraints are set between the inner circumferential surface of the main rubber layer and the outer circumferential surface of the metal inner tube, as well as between the outer circumferential surface of the main rubber layer and the inner circumferential surface of the metal outer tube. This bonding constraint means that in all subsequent load steps, the nodes on the bonded interface always maintain displacement coordination, that is, there is no relative motion degree of freedom, thus truly reflecting the common load-bearing characteristics of the vulcanized integrated structure.

[0033] The second type involves face-to-face mechanical contact between two separate bushings. For the two opposing annular flanged end faces on the two separate bushings, a face-to-face contact pair needs to be established between them. In this contact pair, one flanged end face is defined as the primary contact surface, and the other as the secondary contact surface. The normal contact behavior is set to hard contact, meaning that the contact surfaces are allowed to transmit normal pressure under compression, but not to penetrate each other. Simultaneously, the tangential contact behavior is set to Coulomb friction contact, with a reasonable friction coefficient input to simulate the tangential friction force generated by the relative sliding of the two flanged end faces during assembly and compression. Establishing this face-to-face contact relationship is crucial; it forms the basis for accurately simulating the pre-compression deformation of the flanges and the independent load-bearing capacity of the flanged structure in subsequent steps.

[0034] To discretize the continuous geometric solids into finite element units suitable for numerical solution, the model is meshed. Given that the main rubber layer and its annular flange structure will undergo significant shape changes during subsequent pre-compression and stiffness calculations, and to accurately capture the stress gradient changes in the flange bending region, sufficiently fine solid elements are used for meshing the main rubber layer and the annular flange structure. Hexahedral elements are preferred to ensure good convergence and computational accuracy in large deformation analysis. For the inner and outer metal tubes, due to their minimal deformation and primary role as rigidity transfer components, relatively coarser solid element meshes can be used, balancing the overall computational scale and solution efficiency, while ensuring coordination with the rubber layer mesh nodes.

[0035] Through the above-mentioned progressive steps, a three-dimensional finite element simulation model of the two halves of the assembled bushing, which includes complete geometric shape, accurate material properties, and real connection and contact relationships, can be established. The model has been assigned boundary conditions that reflect its initial manufacturing state, providing a ready numerical platform for the simulation of the subsequent flanged pre-compression assembly process.

[0036] S2. Apply an axial assembly clamping load to the three-dimensional finite element simulation model to drive the two split bushings to move towards each other, causing the two annular flange structures to undergo pre-compression deformation. Solve for the pre-compression stress field and set the pre-compression stress field as the initial stress field.

[0037] Specifically, before performing pre-compression loading, boundary conditions that reflect the actual assembly tooling constraints must first be applied to the model.

[0038] First, a complete fixed constraint is applied to the outer metal tube of one of the split bushings, that is, the outer metal tube is constrained in all translational and rotational degrees of freedom in the axial, radial and circumferential directions, so that it remains stationary during the simulation process, simulating the fixed state of the fixed-side clamping component in the actual assembly tooling.

[0039] Secondly, an axial forced displacement driving load is applied to the outer metal tube of another split bushing. The direction of this driving load is along the bushing axis and points towards the aforementioned fixed split bushing, that is, it drives the split bushing to move axially toward the fixed side. When applying this forced displacement, the radial and circumferential degrees of freedom of the outer metal tube must be constrained simultaneously to ensure that it does not eccentrically offset or rotate during axial movement, retaining only the axial translational degree of freedom.

[0040] In addition, for the metal inner tubes in the two split bushings, no active load is applied during the pre-compression stage. Their motion is entirely driven by the deformation of the main rubber layer and the annular flange structure, in order to simulate the free state when the mandrel has not yet been inserted during the actual assembly process.

[0041] According to the product design specifications, the original thickness of the ring-shaped flange structure in its free state is 5.3 mm. After assembly and clamping, the flange thickness needs to be compressed to 4.0 mm, thus determining the single-sided compression amount to be 1.3 mm.

[0042] Based on this, the total amplitude of the axial forced displacement is set to 1.3 mm. The physical meaning of this displacement is: under the driving load, the metal outer tube of the movable side split bushing moves 1.3 mm axially towards the fixed side split bushing. Since the two annular flange end faces are only separated by a very small assembly gap in the initial modeling, this 1.3 mm opposing displacement will drive the two flange end faces to first eliminate the initial gap, then generate contact compression, and finally cause the flanged rubber structure to produce a total compressive deformation of 1.3 mm, that is, the flange thickness is reduced from 5.3 mm in the free state to 4.0 mm after compression.

[0043] In controlling the loading process, the total displacement of 1.3 mm is divided into multiple uniform increment steps, for example, setting one hundred equally spaced loading substeps, with each substep applying a displacement increment of 0.013 mm. The loading strategy of gradually increasing small step sizes aims to ensure a smooth transition of the contact state during the subsequent nonlinear solution process, and to avoid contact penetration or solution divergence due to excessively large single displacement increments.

[0044] Because the pre-compression process involves the hyperelastic nonlinearity of the rubber material, the geometric nonlinearity caused by the bending deformation of the flange structure, and the state nonlinearity of the two flange end faces from separation to contact and then to mutual compression, an implicit nonlinear finite element solver must be used for quasi-static solution. The calculation mechanism is as follows: Within each increment step, the solver first predicts the displacement field response of the model based on the current displacement boundary conditions, and then calculates the stress tensor generated by deformation within the rubber element. For the main rubber layer and the annular flange structure with a hyperelastic constitutive model, the solver, based on the input hyperelastic material constants, differentiates the current strain state using the strain energy density function and updates the tangent stiffness matrix in real time for that increment step.

[0045] Building upon this, the solver introduces a contact algorithm to handle the face-to-face contact pair between the two annular flange end faces. At the start of each incremental step iteration, the solver first performs a contact search to determine the geometric gap state between the primary and secondary contact surfaces. If the gap is positive, the contact surfaces are in a separated state and no contact force is transmitted; if the gap closes to zero or becomes negative, the normal contact constraint is activated, a hard contact condition is applied to prevent node penetration, and the contact normal pressure is calculated. Simultaneously, based on the set Coulomb friction coefficient, if there is a tangential relative sliding tendency between the contact surfaces, the solver further calculates the tangential friction force and synchronously updates it to the global load vector.

[0046] For each increment step, the solver employs the Newton-Raphson iterative algorithm for equilibrium iteration. In each iteration, the solver calculates the residuals of internal and external forces under the current configuration, i.e., the difference between the externally applied axial driving force and the internal resistance generated by rubber deformation and contact forces. If the norm of this residual vector is greater than a preset convergence tolerance, the solver corrects the displacement increment based on the residuals, recalculates the strain, stress, and contact state, and enters the next iteration cycle. This process is repeated until the residuals of internal and external forces are reduced to within the convergence tolerance, and the contact penetration is also less than the allowable value. Only then can the current increment step be considered to have achieved force and contact equilibrium, and the solver proceeds to the next increment step.

[0047] Through the tightly coupled calculation of incremental loading and iterative balancing described above, the nonlinear solution for the entire pre-compression assembly condition is completed until the total displacement of 1.3 mm is fully applied and the final incremental step converges.

[0048] During the above solution process, the flange structure of the two split bushings underwent a series of continuous physical deformation evolutions. In the initial loading stage, the two opposing annular flange end faces first came into contact, with the initial contact area being only a small local area of ​​the flange end faces. As the axial displacement continued to increase, the contact area gradually expanded from the center of the end faces towards the inner and outer circumferences of the flanges. The cantilevered portion of the flanges began to bend and compress, and the rubber material flowed radially outwards and inwards under the contact pressure. When the displacement approached the 1.3 mm endpoint, the flange thickness was uniformly compressed from 5.3 mm to 4.0 mm, and the contact state of the flange end faces developed from point contact to complete surface contact. The rubber material inside the flanges stored significant compressive strain energy due to triaxial compressive stress. Throughout the process, the main rubber layer only experienced minimal elastic deformation due to the flange deformation, and its stress level was far lower than that of the flange area.

[0049] After the nonlinear solution is completed and the final incremental step converges, the post-processing extraction stage begins. This stage requires extracting two types of key data from the solution results: deformation configuration data and stress field distribution data.

[0050] To extract the deformation configuration data, the spatial coordinate values ​​of all finite element nodes in the last incremental step need to be read from the results file. These coordinate values ​​represent the actual deformed geometry of the flange structure when compressed to a thickness of 4.0 mm. In particular, it is crucial to extract the positional offsets of all nodes in the two annular flange regions to confirm whether the compression in the thickness direction meets the preset target of 1.3 mm.

[0051] To extract stress field distribution data, the stress tensor components at all Gaussian integration points in the last incremental step need to be read from the results file. For hyperelastic rubber materials, the focus is on extracting the components of the Cauchy stress tensor, especially the compressive stress component along the thickness direction and the shear stress component in the flange bending region. Using a nodal stress interpolation algorithm, the stress tensor at the integration points is extrapolated to the element nodes, thereby constructing a continuously distributed pre-compression stress field throughout the flange region.

[0052] S3. Based on the initial stress field, apply a radial displacement load to the inner metal tube, solve for the radial stiffness curve, and lock the structural parameters of the main rubber layer that meet the radial support requirements.

[0053] Specifically, radial stiffness simulation aims to simulate the mechanical response of the bushing when subjected to radial loads perpendicular to the axis. This indicator is directly related to the installation positioning accuracy and support stability of the thermal management module under vehicle driving vibration conditions.

[0054] During radial stiffness simulation, the outer metal tubes of the two split bushings are kept completely fixed to simulate the constrained state after being press-fitted into the support base. Subsequently, a radial displacement load is applied to the inner circumferential surfaces of the two inner metal tubes. This displacement load is applied along a certain radial direction of the bushing, starting from zero and gradually increasing in predetermined increments until the maximum radial design displacement. In this embodiment, the maximum radial displacement is set to 2.0 mm, and the entire loading process is divided into several uniform substeps to ensure solution convergence.

[0055] During loading, an implicit nonlinear solver extracts the resultant radial support reaction force of the inner metal tube at each increment step, generating a complete radial stiffness curve with radial displacement as the abscissa and radial support reaction force as the ordinate. The radial stiffness value is obtained by fitting the linear segment of this curve. In this embodiment, simulation calculations show that within the radial displacement range of 0 to 2.0 mm, the radial support reaction force linearly increases from zero to approximately 1200 Newtons, and the fitted radial stiffness is approximately 600 Newtons per millimeter. Furthermore, the curve exhibits good linearity, indicating that the main rubber layer is within a stable elastic deformation range within this radial displacement interval.

[0056] After obtaining the radial stiffness value, it is compared with the preset target value for the thermal management module support stiffness. If the simulated value falls within the target range, the current main rubber layer structural parameters—including the main rubber wall thickness, axial length, and inner and outer tube diameters—are deemed to meet the radial support requirements. This set of radial structural parameters is then locked and used as fixed parameters in all subsequent axial optimization analyses. It is important to emphasize that "locking" here means that in subsequent axial segmented simulations and parameter iterative optimizations, no further modifications are made to the geometry of the main rubber layer; only the relevant parameters of the annular flange structure are adjusted. This "radial locking" strategy fundamentally severs the parameter coupling path between radial and axial stiffness.

[0057] S4. After locking the structural parameters of the main rubber layer, apply axial displacement load continuously along the axial direction, and divide the entire axial loading stroke into the first axial displacement interval and the second axial displacement interval with the preset displacement threshold as the dividing point. Extract the axial displacement-reaction data of each interval to obtain the segmented axial stiffness curve.

[0058] Specifically, the goal of axial stiffness simulation is not to obtain a single axial force-displacement curve, but to divide the entire axial travel into two independent displacement intervals based on differences in physical load-bearing mechanisms, and extract the stiffness characteristics of each interval separately. This method adopts a "continuous loading, segmented extraction" strategy: during a continuous axial displacement loading process, the loading process is logically divided into two intervals based on a displacement threshold of 1.3 mm. In the post-processing stage, the displacement-reaction data for each interval are extracted separately, and their respective stiffness values ​​are calculated. This strategy ensures both the physical continuity of the loading process and the interval independence of stiffness evaluation.

[0059] It needs to be clarified again that the single-sided pre-compression of the flange in step two is exactly 1.3 mm, which is consistent with the upper limit of the first axial displacement range. This "consistency" is not a coincidence, but one of the design logics of the present invention: the flange is pre-compressed by 1.3 mm during assembly, which means that in the subsequent small axial displacement buffer section (i.e., an additional 0 to 1.3 mm of axial tension or compression is applied), the flange continues to deform or partially unload from the pre-compressed state. Its stiffness response is entirely dominated by the bending and compression characteristics of the flange structure itself, while the main rubber layer hardly participates in the load-bearing, thereby achieving a low-stiffness buffering effect.

[0060] After completing the radial simulation and locking the radial parameters, restart the analysis or continue the loading step based on the pre-compression initial stress field obtained in step two. At this point, no active load is applied to the outer metal tube, keeping it in a fixed state. An axial displacement load is applied to the common end face of the two inner metal tubes. This displacement direction is opposite to the compression direction during pre-compression, simulating the bushing's axial tension or unloading rebound during operation. The displacement load starts from zero and gradually increases to 1.3 mm.

[0061] During the displacement increment of 0 to 1.3 mm, the mechanical response of the flange transitioned from a pre-compression state to unloading or reverse loading. Since the flange was compressed to a thickness of 4.0 mm during assembly, when the inner tube was subjected to axial tensile displacement, the contact pressure between the flange end faces gradually decreased, the flange rubber elastically recovered, and the bending area of ​​the flange underwent bending deformation. Within this displacement range, the annular flange structure is the main load-bearing and deformation component; the bending stiffness and local compressive stiffness of its rubber material together determine the axial reaction force level of this section. Because the main rubber layer is located between the inner and outer metal tubes and its axial length remains constant, it only experiences minimal tensile deformation within this small displacement range, and its contribution to the axial reaction force is negligible.

[0062] The solver incrementally records the axial displacement and corresponding axial reaction force until the displacement reaches 1.3 mm. Post-processing extracts the axial reaction force value at the endpoint of this interval, i.e., the 1.3 mm displacement, and calculates the average stiffness within the 0 to 1.3 mm range. In this embodiment, simulation shows that at a 1.3 mm axial displacement, the axial reaction force is approximately 155 Newtons, and the average stiffness is approximately 119 Newtons per millimeter. This low stiffness value meets the buffer design requirements under small-amplitude vibrations and can effectively attenuate the transmission of high-frequency micro-amplitude vibrations to the thermal management module.

[0063] The loading in the second interval does not start from zero again, but rather continues to apply axial displacement load in the same direction based on the end state of the first interval, increasing it from 1.3 mm to 3.3 mm, that is, the incremental stroke of this interval is 2.0 mm.

[0064] When the axial displacement exceeds 1.3 mm, the state of the flange structure undergoes a fundamental change. The flange, compressed to 4.0 mm during the pre-compression stage, gradually recovers its thickness under reverse tension, and the contact pressure between the two flange end faces continuously decreases until they completely separate. Once the flange end faces separate, the annular flange structure no longer bears the compressive load, and its further deformation capacity is extremely limited. At this point, the transmission path of the axial load changes—the load is no longer borne by the contact compression of the flange, but is directly transmitted to the main rubber layer through the inner metal tube, and then the axial shear and compressive deformation of the main rubber layer bears the subsequent displacement load. Because the axial stiffness of the main rubber layer is much higher than the bending stiffness of the flange structure, the rate of increase in axial reaction force with displacement in this range increases significantly.

[0065] The solver continues to incrementally record axial displacement and axial reaction force data within the interval from 1.3 mm to 3.3 mm. Post-processing extracts the axial reaction force values ​​at the start (1.3 mm) and end (3.3 mm) of this interval and calculates the segmented stiffness of this interval using the formula: segmented stiffness equals (reaction force at 3.3 mm minus reaction force at 1.3 mm) divided by (3.3 mm minus 1.3 mm). In this embodiment, simulation shows that the axial reaction force at the 3.3 mm displacement is approximately 715 Newtons. Subtracting the 155 Newtons at 1.3 mm, the difference of 560 Newtons is divided by the incremental stroke of 2.0 mm, resulting in a segmented stiffness of approximately 280 Newtons per millimeter. This high and stable stiffness provides reliable limiting support under large displacement conditions, preventing excessive movement of the thermal management module.

[0066] S5. Using the original thickness of the annular flange structure as the iterative variable, establish multiple simulation models with different original flange thicknesses. Repeat the pre-compression simulation and segmented stiffness solution. Compare the segmented axial stiffness curves obtained from each simulation with the preset design target curve, select the optimal original flange thickness, and calculate the corresponding assembly pre-compression amount in reverse.

[0067] Specifically, the design target curve is not an arbitrarily given mathematical function, but rather originates from the system decomposition of the vehicle-level performance target to the bushing component level. Its establishment process has strict engineering logic and specifically includes the following three progressive levels.

[0068] The first level involves establishing overall vehicle NVH performance targets. In the early stages of vehicle engineering development, OEMs set limits for in-vehicle vibration and noise under idling, normal cruising speed, and rapid acceleration / deceleration conditions, based on model positioning and market requirements. These limits are further broken down to various vibration isolation components such as the powertrain mounting system, subframe bushings, and thermal management module mounting points, ultimately forming requirements for vibration isolation rates and displacement control at the thermal management module mounting points.

[0069] The second level involves the transformation of the dynamic requirements of the bushing mounting points. Based on the mass, center of mass location, and mounting point layout of the thermal management module, the dynamic force amplitude and maximum allowable displacement at each mounting point under typical operating conditions are calculated. Under small-amplitude vibration conditions, the bushing is typically required to have low dynamic stiffness to attenuate high-frequency vibration transmission, corresponding to an axial displacement range of approximately 0 to 1.3 mm. The upper limit of stiffness in this range is determined by the vibration isolation rate requirement, while the lower limit is constrained by the module positioning accuracy requirement. Under large-displacement conditions, such as when a vehicle travels over bumpy roads or brakes suddenly, the bushing is required to provide sufficient and stable support within a displacement range of 1.3 to 3.3 mm to limit excessive movement of the module. The lower limit of stiffness in this range is determined by the limiting capacity requirement, while the upper limit is constrained by the strength bearing limit of the mounting bracket.

[0070] The third level involves the mathematical expression and correction of the design target curve. The stiffness requirements of the two intervals mentioned above are transformed into several key control points on the axial displacement-reaction target curve. The first interval starts at the origin and ends at the target reaction force corresponding to a 1.3 mm displacement; the line connecting these two points represents the target stiffness slope for this interval, calculated based on the vibration isolation rate. The second interval starts at the target reaction force at a 1.3 mm displacement and ends at the target reaction force corresponding to a 3.3 mm displacement; the target stiffness slope for this interval is calculated based on the limiting capacity. Furthermore, a smooth transition requirement needs to be set at the 1.3 mm dividing point, meaning the first derivative of the curve should not show abrupt changes at this point to avoid the impact of stiffness jumps. After obtaining these control points, designers fine-tune and correct them using benchmark data from similar vehicle models and historical project experience, ultimately forming a complete design target curve containing the coordinates of key control points, segmented target slopes, and smoothness requirements for transition points, serving as a benchmark for subsequent scheme comparisons.

[0071] After defining the target design curve, the next step is to select the iterative variables. In this method, the original thickness of the flange is selected as the main iterative variable. The engineering basis for this selection is that the flange thickness is the most sensitive parameter determining the bending stiffness of the flange structure. Without changing any dimensions of the main rubber layer, adjusting only the original thickness of the flange can significantly change its resistance level during compression, thereby directly affecting the low stiffness value in the first interval. At the same time, the flange thickness also determines its density after compression to the limit, thus indirectly affecting the starting point and transition rate of load transfer to the main rubber layer.

[0072] Based on the above analysis, and building upon the baseline geometric model established in step one, all dimensional parameters of the inner metal tube, outer metal tube, and main rubber layer, as well as the properties of the rubber material, remain completely unchanged. Only the original thickness parameter of the annular flange is altered. In this embodiment, using the initially designed 5.3 mm as the center value, three deviation values ​​are taken on either side: 5.0 mm, 5.3 mm, and 5.6 mm. Each set of schemes requires a reconstructed geometric model and a complete mesh generation.

[0073] The selection logic for the three schemes is as follows: the 5.0 mm scheme represents a thinner flange thickness with lower bending stiffness, expected to generate lower axial reaction force in the first interval. Simultaneously, the flange is more easily compressed to a dense state after pre-compression, potentially triggering the main rubber bearing capacity earlier. The 5.6 mm scheme represents a thicker flange with higher bending stiffness, expected to generate higher axial reaction force in the first interval, with a reduced buffering effect, but stronger durable bearing capacity. The 5.3 mm scheme serves as an intermediate value, aiming to achieve a balance between the two. These three schemes cover three typical positions in the design space—upper, middle, and lower—effectively determining the trend of the objective function within the variable range.

[0074] After establishing three sets of geometric models with different flange thicknesses, all operations of steps two and three are performed independently and completely identically for each set of models.

[0075] Specifically, for the design with a flange thickness of 5.0 mm, the pre-compression assembly simulation in step two is first performed, compressing the flange from 5.0 mm to the set target assembly thickness of 4.0 mm, at which point the compression on one side is 1.0 mm. Then, using this pre-compression state as the initial condition, the axial segmented loading in step three is performed, extracting displacement-reaction curve data for the two intervals of 0 to 1.3 mm and 1.3 to 3.3 mm respectively, to obtain the two segments of axial stiffness characteristics of the design.

[0076] For the scheme with a flange thickness of 5.3 mm, the pre-compression process compresses the flange from 5.3 mm to 4.0 mm, with a single-sided compression of 1.3 mm. This scheme is the one for which simulation has been completed and initial results obtained in step three, and its data can be directly reused without repeated calculations.

[0077] For the design with a flange thickness of 5.6 mm, a pre-compression assembly simulation is first performed to compress the flange from 5.6 mm to 4.0 mm, at which point the compression on one side is 1.6 mm. After pre-compression, axial segmented loading is also performed, and stiffness curve data for the two intervals are extracted.

[0078] It is worth noting that the flange thickness after assembly was uniformly set to 4.0 mm in all three schemes. This means that schemes with thicker flanges require greater clamping force during assembly, and the pre-compression stress stored inside the flange after assembly is also higher. This uniform thickness setting ensures that the comparison of the three schemes is based on the same assembly boundary conditions, guaranteeing the fairness of the comparison results.

[0079] After obtaining all simulation data for the three schemes, the comparison and judgment phase begins. In this phase, the simulated piecewise stiffness curves of each scheme are compared with the design target curves established in the first part of step four. The comparison is carried out one by one based on the following three parallel judgment criteria.

[0080] The first criterion is the stiffness matching degree of the first interval. The slope of the simulated curve in the range of 0 to 1.3 mm is compared with that of the target curve. The average stiffness value of the simulated curve should fall within the error band of ±5% relative to the target curve. If the simulated stiffness is too high, it indicates that the flange thickness is too large and the buffering effect is insufficient; if the simulated stiffness is too low, it indicates that the flange thickness is too thin, which may lead to insufficient module positioning stiffness.

[0081] The second criterion is the stiffness matching degree of the second interval. The average slope of the simulated curve in the 1.3 to 3.3 mm interval is compared with the target curve, and the error is required to be controlled within ±5%. The stiffness in this interval is mainly contributed by the main rubber layer, but because different flange thicknesses affect the load transfer rate, the stiffness in this interval may also differ between different schemes.

[0082] The third criterion is the continuity of the transition point. It requires that at the dividing point where the axial displacement equals 1.3 mm, the two segments of the simulation curve maintain a mechanically continuous and smooth transition. Specifically, the axial reaction force at this point should not exhibit a sudden jump, and the rate of change of the tangent slope of the left and right segments of the curve at this point should be controlled within a reasonable range. If a significant stiffness step occurs at the transition point, it will lead to an impact sensation near that displacement, degrading NVH performance.

[0083] When making the above three judgments, the comparison results of the three schemes are as follows: The scheme with a flange thickness of 5.0 mm has low stiffness in the first interval, and the reaction force increases slowly, failing to meet the lower limit requirement of the buffer stiffness of the target curve. At the same time, the reaction force at 1.3 mm is low, causing the starting point of the second interval to deviate from the target, and the first criterion is not met. The scheme with a flange thickness of 5.6 mm has significantly high stiffness in the first interval, and the buffer section is too rigid, losing the ability to attenuate high-frequency micro-amplitude vibrations, and also fails to meet the first criterion. Only the scheme with a flange thickness of 5.3 mm has its average stiffness in the first interval, average stiffness in the second interval, and reaction force value at the 1.3 mm transition point all within the error band of the target curve, and the transition is continuous and smooth. Therefore, 5.3 mm is determined to be the only optimal original flange thickness parameter among the three schemes.

[0084] After locking the original flange thickness at 5.3 mm, reverse calculations are required to confirm the assembly compression amount and related process parameters. Here, "reverse" refers to the process relative to the forward design process—the forward process predicts stiffness performance from known assembly parameters, while reverse calculation starts from a confirmed solution that meets the stiffness target and works backward to confirm the physical quantities that need to be controlled in the assembly process.

[0085] As mentioned earlier, the original thickness of the flange is 5.3 mm, and the target thickness of the flange after assembly is 4.0 mm. Therefore, the single-sided compression is directly calculated as the difference between the two, which is 1.3 mm. This 1.3 mm compression is exactly consistent with the upper limit of the first axial displacement range, which verifies the matching relationship between the flange pre-compression stroke and the small displacement buffer stroke from a numerical perspective.

[0086] Furthermore, the magnitude of the axial clamping force applied to the outer metal tube when the flange is compressed from 5.3 mm to 4.0 mm is extracted from the pre-compression simulation results of step two. This clamping force is a key input parameter for the assembly tooling design, including the output range of the pressing equipment and the criteria for determining the end point of the pressing stroke. Simultaneously, based on the stress field distribution of the flange in the pre-compression simulation, the average contact pressure on the flange end face can be calculated. This parameter is used to verify the creep and relaxation risk of the flange rubber under long-term compression.

[0087] In addition, it is necessary to reverse-check whether the compression exceeds the allowable compression limit of the rubber material. According to the mechanical properties of hyperelastic materials, the stress-strain curve of rubber exhibits obvious "stiffening" after the compressive strain reaches a certain level. If the compression is too large, it will lead to stress concentration and even damage inside the flange. In this step, by extracting the maximum principal strain value of the flange area in the pre-compression simulation and comparing it with the material's ultimate failure strain, it is confirmed that the compression of 1.3 mm is within the safety margin range.

[0088] Finally, the reverse calculation outputs three sets of interrelated key process parameters: original flange thickness of 5.3 mm, flange thickness after assembly of 4.0 mm, single-sided pre-compression of 1.3 mm, and reference value for assembly clamping force. These parameters constitute a complete and self-consistent assembly parameter system, which can be directly transferred to step five for mold design and tooling development.

[0089] S6. Output the finalized structural parameters, including the optimal original flange thickness and assembly pre-compression amount, to the process end.

[0090] Specifically, the optimal structural parameters confirmed through iterative optimization are organized and finalized to form a complete process parameter package. Specific output parameters include: the axial length of a single split bushing is 15 mm, the rubber thickness in the free state of the flange is 5.3 mm, the compressed thickness of the flange after assembly is 4.0 mm, the single-sided compression during assembly clamping is 1.3 mm, and the assembly clamping force range obtained from pre-compression simulation.

[0091] The aforementioned finalized parameters are directly output to the process end for designing the cavity dimensions of the vulcanizing mold and setting the clamping stroke and clamping force of the assembly clamping fixture. Since all key parameters have been determined through simulation iteration, mold development and tooling manufacturing can be completed in one go, without the need for multiple rounds of mold trials and repairs.

[0092] After the actual tooling prototype was fabricated, it was installed on a test bench for stiffness testing. Comparison of the measured results with the simulation prediction curves showed that the trends were consistent, the segmented stiffness characteristics were well matched, and the error was within 5%, verifying the effectiveness and reliability of the simulation method.

[0093] Other embodiments of this application will readily occur to those skilled in the art upon consideration of the specification and practice of the disclosure herein. This application is intended to cover any variations, uses, or adaptations of this application that follow the general principles of this application and include common knowledge or customary techniques in the art not disclosed herein.

[0094] It should be understood that this application is not limited to the precise structure described above and shown in the accompanying drawings, and various modifications and changes can be made without departing from its scope.

Claims

1. A CAE segmented stiffness design method for thermal management bushings of flanged pre-compression and mating structures, characterized in that, Includes the following steps: A three-dimensional finite element simulation model of a two-part assembled split bushing is established. The split bushing includes an inner metal tube, an outer metal tube, a main rubber layer located between the two, and an annular flange structure set at the axial end of the main rubber layer. An axial assembly clamping load is applied to the three-dimensional finite element simulation model to drive the two split bushings to move towards each other, causing the two annular flange structures to undergo pre-compression deformation. The pre-compression stress field is obtained by solving the problem and is set as the initial stress field. Based on the initial stress field, a radial displacement load is applied to the inner metal tube, the radial stiffness curve is obtained by solving, and the structural parameters of the main rubber layer that meet the radial support requirements are locked accordingly. After locking the structural parameters of the main rubber layer, axial displacement load is continuously applied along the axial direction, and the entire axial loading stroke is divided into the first axial displacement interval and the second axial displacement interval with the preset displacement threshold as the dividing point. The axial displacement-reaction data of each interval are extracted to obtain the segmented axial stiffness curve. Using the original thickness of the annular flange structure as the iterative variable, multiple simulation models with different original flange thicknesses are established. The pre-compression simulation and segmented stiffness solution are repeatedly executed. The segmented axial stiffness curves obtained from each simulation are compared with the preset design target curves to select the optimal original flange thickness. The corresponding assembly pre-compression amount is then calculated in reverse. The finalized structural parameters, including the optimal original flange thickness and assembly pre-compression, are output to the process end.

2. The CAE segmented stiffness design method for the thermal management bushing of the flanged pre-compression and mating structure as described in claim 1, characterized in that, When establishing a three-dimensional finite element simulation model, two axially symmetrical and structurally identical split bushing geometric entities are constructed. The two split bushings are arranged in the same global coordinate system in a coaxial alignment manner, and their annular flange end faces each other in the axial direction, forming an overall structural form of two halves assembled together.

3. The CAE segmented stiffness design method for the thermal management bushing of the flanged pre-compression and mating structure as described in claim 1, characterized in that, The inner and outer metal tubes adopt a linear elastic constitutive model, and their respective elastic moduli and Poisson's ratios are input. The main rubber layer and the annular flange structure adopt a hyperelastic constitutive model. The hyperelastic material constants are determined by fitting the mechanical test data of the standard rubber specimen, and the main rubber layer and the annular flange structure are given the same hyperelastic material parameters.

4. The CAE segmented stiffness design method for the thermal management bushing of the flanged pre-compression and mating structure as described in claim 1, characterized in that, The inner circumferential surface of the main rubber layer and the outer circumferential surface of the metal inner tube, as well as the outer circumferential surface of the main rubber layer and the inner circumferential surface of the metal outer tube, are both set as adhesive binding constraints to simulate the integral fixed connection state of vulcanization bonding; the annular flange end faces that are opposite each other on the two split bushings are set as face-to-face contact pairs, and normal hard contact behavior and tangential Coulomb friction contact behavior are set to simulate the assembly contact and relative sliding state when the two halves of the bushing are assembled.

5. The CAE segmented stiffness design method for the thermal management bushing of the flanged pre-compression and mating structure as described in claim 1, characterized in that, When applying an axial assembly clamping load to simulate the pre-compression of the flange, a completely fixed constraint is applied to the outer metal tube of one split bushing, and an axial forced displacement driving load is applied to the outer metal tube of the other split bushing, causing the two split bushings to move towards each other along the axial direction until the thickness of the annular flange structure is compressed from the free state to the target assembly thickness. After the solution is completed, the deformation configuration and stress field distribution data of the flange region under the target compression thickness are extracted, and the stress field is written into the initial stress field of the subsequent stiffness calculation.

6. The CAE segmented stiffness design method for the thermal management bushing of the flanged pre-compression and mating structure as described in claim 5, characterized in that, The value of the preset displacement threshold is equal to the value of the single-sided pre-compression of the annular flange structure during the assembly process.

7. The CAE segmented stiffness design method for the thermal management bushing of the flanged pre-compression and mating structure as described in claim 5, characterized in that, When performing axial segmented stiffness simulation, in the first axial displacement interval from zero position to a preset displacement threshold, the axial reaction force is contributed by the bending deformation and compression deformation of the annular flange structure. The average stiffness of this interval is extracted to match the small-amplitude vibration buffer target. In the second axial displacement interval from the preset displacement threshold to the maximum axial displacement, the two annular flange end faces are out of contact and compression state, and the axial load is mainly borne by the main rubber layer. The average stiffness of this interval is extracted to match the set displacement limit support target.

8. The CAE segmented stiffness design method for the thermal management bushing of the flanged pre-compression and mating structure as described in claim 7, characterized in that, Using the original thickness of the annular flange structure as the iterative variable, multiple simulation models with different original flange thicknesses are established, including: Using the original thickness of the annular flange as the sole iterative variable, while keeping all dimensional parameters and material properties of the inner metal tube, outer metal tube, and main rubber layer unchanged, a deviation value is taken on each side of the preset center value to form at least three simulation schemes with different original flange thicknesses. The target compression thickness of the flange after assembly remains consistent in each scheme.

9. The CAE segmented stiffness design method for the thermal management bushing of the flanged pre-compression and mating structure as described in claim 8, characterized in that, When comparing and screening the segmented axial stiffness curves obtained from each group of simulations with the preset design target curves, a comprehensive evaluation is made based on the following three criteria: whether the error between the average stiffness of the simulation curve and the target curve in the first axial displacement interval falls within the preset allowable range; whether the error between the average stiffness of the simulation curve and the target curve in the second axial displacement interval falls within the preset allowable range; and whether the stiffness transition of the simulation curve at the preset displacement threshold boundary point meets the continuity requirement.

10. The CAE segmented stiffness design method for the thermal management bushing of the flanged pre-compression and mating structure as described in claim 9, characterized in that, After determining the optimal original flange thickness, when calculating the assembly pre-compression amount in reverse, the target compression thickness after assembly is subtracted from the optimal original flange thickness to obtain the single-sided pre-compression amount; at the same time, the assembly clamping force corresponding to the compression amount is extracted from the pre-compression simulation results. The original flange thickness, the target compression thickness after assembly, the single-sided pre-compression amount, and the assembly clamping force are all output as structural parameters to the process end.

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