Selecting materials for optimum performance.
Engineering plastics are now entering markets where their mechanical performance must meet increasingly demanding requirements, yet companies cannot afford over-designed parts or lengthy, iterative product development cycles. Therefore, engineers must have design technologies that allow them to productively create the most cost-effective design with the optimum selection of materials and process.
The design engineering process involves meeting end-use requirements with the lowest cost design, material, and process combination. Design activities include creating geometries and performing engineering analysis to predict part performance. Material characterization provides engineering design data, and process selection includes process/design interaction knowledge. In general, the challenge in designing with structural plastics is to develop an understanding not only of design techniques, but also of processing and material behavior.
Engineering thermoplastics exhibit complex behavior when subjected to mechanical loads. Standard data sheets provide overly simplified, single-point data and are either ignored or, if used, probably misleading. The objective of this article is to suggest an alternative that will allow engineers to select an appropriate material.
Over the past few years many plastics databases have been introduced. While many of these databases offer "user-friendly" computer interfaces and sophisticated search programs, they are still primarily "electronic data sheets," lacking procedures to predict part performance. Some databases provide engineering data(1) over a range of application conditions, and knowledge-based materials selection programs have been written(2). However, the methodology for optimum selection of materials and process conditions to meet part performance needs has not been widely used. This methodology will be clarified by descriptions and demonstrations of simple tools and techniques for the initial prediction of part performance.
There is a wide variety of part performance requirements. Some, such as flammability; transparency; ultraviolet stability; electrical, moisture, and chemical compatibility; and agency approvals, are specified as absolute values or simplified choices. However, mechanical requirements such as stiffness, strength, impact, and temperature resistance cannot be specified as absolute values. For example, a part may be required to have a certain stiffness, i.e., maximum deflection for a given loading condition. The part geometry (design) and the material stiffness combine to produce the part stiffness. Thus, it is impossible to select a material without some knowledge of the part design. Similarly, the part may be required to survive a certain drop test or a certain temperature/time/loading condition. Again, it is impossible to select a material by using traditional, inadequate, single-point data such as notched Izod or HDT. In addition, it is important to consider the effects of the design and materials selection of a part on its fabrication. Considerations such as flow and cycle time should be quantitatively included in the material selection process. Simple yet extremely useful tools and techniques for the initial prediction of part performance are presented here.
Many thermoplastic parts are plate-like structures that can be treated as a simply supported plate, possibly reinforced with ribs. An interactive menu-driven computer program, Ribstiff, was developed to provide quick, approximate solutions for the stiffness of laterally loaded rib-stiffened plates(3). The program employs the Rayleigh-Ritz energy method, and is capable of including the geometric nonlinearities associated with the large-displacement response typical of low-modulus materials such as thermoplastics. The user inputs the important parameters of specific plate structures (length, width, thickness, number of ribs, rib geometry), the boundary conditions (simply supported, clamped, point supported), and the loading (central point, uniform pressure, or torsion loading). With the capability of multiple rib pattern definitions, the user can then plot the load-deflection curves for different designs to select the one that is most effective for the specific application. This entire process takes a few minutes--a significant improvement over the time required for a finite-element analysis. This tool has been validated with finite-element results. Figure 2 is an example of Ribstiff's ability to predict the nonlinear load-displacement response.
Part Strength and Impact Resistance
Impact-resistance tests should not only measure the amount of energy absorbed, but also determine the effects of temperature changes on energy absorption. Additionally, they should be able to identify strain-rate-dependent transitions from ductile to brittle behavior and be applicable to a wide variety of geometric configurations. Unfortunately, the test methods available, such as Izod, Gardner, and Dynatup, provide only geometry-specific, single-point data for a specific temperature and strain rate. Also, each test will provide a different value for the ductile/brittle transition. Energy absorption, however measured, is made up of many complex processes involving elastic and plastic deformation, notch sensitivity, and fracture processes of crack initiation and propagation.
The prediction of strength and impact resistance of plastic parts is probably the most difficult challenge for the design engineer. Tensile stress-strain measurements as a function of temperature and strain rate provide one piece of useful information. Most unfilled engineering thermoplastics exhibit ductile behavior in these tensile tests, with increasing strength (maximum stress) as displacement rate increases and/or temperature decreases. However, stress-state effects must be added to the tensile behavior since the three-dimensional stress state created by notches, radii, holes, thick sections, etc., increases the potential for brittle failure. An attempt to include temperature, rate, and stress state is illustrated in the "fracture map" shown in Fig. 3. The results indicate the "strength"' (computed from the failure load per unit thickness) of a notched beam with a notch-root radius of 0.25 mm. The test results are described in detail in Reference 4. The two lines starting in the upper left portion of Fig. 3 increase with increasing loading rate, exhibiting ductile behavior. The increasing failure load with increasing rate and decreasing temperature is consistent with ductile strength measurements from tensile stress-strain tests. The notched Izod result is noted as the single-point value: specific notched beam geometry, room temperature, displacement rate of 3.35 m/sec. The decreasing portions of the two curves are due to brittle behavior promoted by high strain rates and low temperature or the increased plane-strain constraint in the 6.35-mm-thick specimens.
Heat Resistance--Time/Temperature Performance
Polymers exhibit time-dependent deformation (creep and stress relaxation) when subjected to loads. This deformation is significant in many polymers, even at room temperature, and is rapidly accelerated by small increases in temperature. Hence, the phenomenon is the source of many design problems. There is a need for the development and application of methods of predicting whether a component will sustain the required service life when subjected to loading, as the part's useful life could be terminated by excessive deformation or even rupture. For most practical applications of polymers, predictive methods must account for part geometry, loading, and material behavior.
A common measure of heat resistance is the heat-distortion temperature (HDT). A bending specimen 127 by 12.7 mm, with a thickness ranging from 3.2 to 12.7 mm, is placed on supports 102 mm apart, and a load producing an outer fiber stress of 0.46 or 1.82 MPa (66 or 264 psi) is applied. The temperature in the chamber is increased at a rate of 2 |degrees~ C/min. The temperature at which the bar deflects an additional 0.25 mm is called the HDT or, sometimes, the deflection temperature under load (DTUL). Such a test, which involves variable temperature and arbitrary stress and deflection, is of no use in predicting the structural performance of a thermoplastic at any temperature, stress, or time. In addition, it can be misleading when materials are being compared. A material with a higher HDT than another material could exhibit more creep at a lower temperature. Also, some semicrystalline materials exhibit very different values of HDT at 0.46 and 1.82 MPa. For example, HDTs for polybutylene terephthalate (PBT), are 154 |degrees~ C at 0.46 MPa and 54 |degrees~ C at 1.82 MPa. The question that naturally arises is, which HDT to use for comparison with another material that has the same HDT for both stress levels? Another method often used to account for the change in material modulus with temperature is dynamic mechanical analysis (DMA)(5). Although this approach provides a more useful indication of instantaneous modulus variation with temperature than HDT, it is unable to account for the time-dependent nature of most applications. For purposes of predicting part performance and for material selection, tensile creep data are desired.
To be useful for preliminary part design and material selection, creep data must be converted to simple, useful information such as "reduced stiffness" graphs. The applied stress, when divided by the time-dependent strain, produces a time- and temperature-dependent material stiffness. When this "creep modulus" is divided by the room temperature elastic modulus, a "reduced stiffness" is computed. Lines of constant reduced stiffness can be displayed on a temperature-time graph to produce useful information for assessing the heat resistance of a particular material at the time and temperature of the particular application. For example, the results in Fig. 4 for polycarbonate (PC) and a PC/ABS blend indicate that for 10 hrs at 38 |degrees~ C, the two materials would have only a slightly reduced stiffness, while for 1000 hrs at 79 |degrees~ C, the stiffness of the PC/ABS blend would be only one-fourth that of PC and only 10% of its room temperature modulus. This material stiffness must be combined with the part geometry to assess the adequacy of the selected material and design.
Note that for these two materials, Fig. 4 shows the HDT to be consistent with the creep data. For the HDT test, the "equivalent stiffness" required for these materials with the specified displacement (0.25 mm) and stress (1.82 MPa) is computed to be 690 MPa, about 30% of the room temperature stiffness. The 30% reduced stiffness curves reach the HDT at times on the order of the times of the HDT test where significant deformation occurs--above 66 |degrees~ C at 0.5 hr.
The ability to manufacture plastic parts by injection molding is governed by the material behavior, part geometry, and processing conditions. Diskflow(6) is a generic tool capable of analyzing radial flow and quantifying effects of material, geometry, or process changes. Diskflow is composed of a numerical flow analysis, automatic mesh generator, and menu-driven pre- and postprocessors. No knowledge of simulation techniques is required, though a knowledge of injection molding is needed when one is interpreting the results.
Diskflow utilizes modeling techniques common to most commercial analyses, yet it is much faster because of the radial-flow assumption and subsequent numerical methods. Modeling, postprocessing, and CPU times have been minimized, reducing total analysis time from days to minutes. This analysis cannot replace three-dimensional filling analyses as it does not yield any information regarding knit-line and gas-trap locations, cavity pressure and temperature distributions, or more complex mold geometries, but it does yield significant results regarding design feasibility, material performance, and process effects.
Diskflow uses a modified Cross model with (when available) WLF temperature dependence to fit complete viscosity vs. shear rate and temperature data. If this advanced model is not available, the Arrhenius model is employed. Mold and melt temperature can be chosen if default values are not adequate. The mold geometry is defined in terms of a nominal wall thickness and cavity radius that may be calculated by entering cavity volume or projected area. The sprue is defined by entering the sprue type (hot or cold) and then entering the length, upper diameter, and lower diameter.
For flow-length estimation, a constant initial flow rate is assumed, subject to some user-specified maximum pressure limit that mimics a molding machine's capability. As the mold fills at a constant volumetric flow rate the injection pressure will rise as a result of the increasing flow resistance. When the injection pressure attains the user-specified maximum, the analysis switches over to a second phase where the injection pressure is now maintained at a constant value and the flow rate is allowed to vary; the flow rate will eventually decay to zero, at which point a final flow length is attained. The flow length may be defined as the farthest distance that a polymeric material will travel in a mold of some nominal wall thickness given a set of processing conditions. The flow-length capability will examine the feasibility of manufacturing a desired design: If the distance from the gate to the corner of the part is greater than the predicted flow length, then the part may not be manufacturable. Figure 5 shows the dependence of flow length on wall thickness for a maximum injection pressure of 103.5 MPa for polycarbonate. This information is particularly useful in the early stages of design and material selection.
Cycle Time Estimate From Mold-Cooling Analysis
In injection molding, the injected molten thermoplastic is held in the cooled mold cavity until the part solidifies as a result of heat transfer. The time for the part to solidify to the extent that it can be removed from the mold without damage generally is the majority of the total cycle time. The large impact of the cooling time on the total processing cost is obvious.
During the cooling phase, heat conduction is the prime mechanism of heat transfer. The development of a simplified mold-cooling program allows designers and molders to evaluate materials and process parameters in a rapid, convenient, and cost-efficient manner. Plastic parts are usually thin, and thus one-dimensional, transient heat-conduction analysis is adequate to approximate the cooling of the real part. The main assumption is that the mold surface is kept at a constant temperature throughout the cooling phase. Comparing calculated minimum cooling times for different material part geometries (i.e., thickness) and processing conditions can help optimize the material selection process.
Thermal material properties are strong functions of temperature. Because the thermoplastic material experiences a wide range of temperatures during the cooling phase, temperature-dependent material data such as specific heat and thermal conductivity are used for the computations. To perform the analysis the injection temperature, mold temperature, ejection temperature, material, and thickness must be chosen. The program uses a one-dimensional finite-difference scheme to calculate temperature through the thickness as a function of time. When the center of the plate reaches the specified ejection temperature, the analysis is stopped and the results are displayed graphically. If the analysis is performed for a range of part thicknesses, cooling time curves can be produced. These curves can then be used to estimate cycle times in the early stages of material selection and design.
Example: Rib Stiffening vs. Thicker Wall
A simple example is presented to illustrate the preliminary design/material selection process. A 254-mm x 254-mm simply supported plate is loaded at room temperature with a uniform pressure of 760 Pa. The maximum allowable deflection is 3.2 mm. By use of the Ribstiff program, the nonlinear load-displacement response of the plate can be computed. Through iteration, a polycarbonate plate with a thickness of 2.5 mm was determined to satisfy the requirements. From Fig. 5, the flow length is 320 mm. Thus, the plate could be filled with a center gate or from the center of an edge. From Fig. 6, the in-mold cooling time is 10 sec. The plate volume is 0.00016 |m.sup.3~.
A second design can be produced by designing a rib-stiffened plate. Again, through iteration, a 1.5-mm-thick plate with ten ribs in each direction with a rib height of 4.5 mm and a rib thickness of 1.5 mm would meet the deflection requirement. From Fig. 5, the flow length is about 175 mm. Thus, since a centergated plate would have a flow length of 175 mm, the part would probably fill if the ribs served as flow leaders to aid the flow. However, it is generally not advisable to push an injection molding machine to its limits because this will exaggerate inconsistencies in the material and the process. A more thorough three-dimensional process simulation should be performed to determine the viability of this design before it is chosen. From Fig. 6, the in-mold cooling time is about 4 sec, a considerable savings (6 sec/part) in cycle time as compared to the plate with no ribs. In addition, the volume of the ribbed plate is 0.00013 |mm.sup.3~, a saving of 20% on material compared with the plate with no ribs. Since the ribs would produce a constrained, three-dimensional stress state, consideration of impact would be important for high rates of loading and low temperature. The fracture map shows a tendency for brittle behavior with polycarbonate at low temperature and high loading rates for notched or constrained geometries.
If time/temperature performance were added to this example as a requirement, the optimum material may change or the preliminary design would need to be modified. If the same load were applied to the plate for 1000 hrs at a temperature of 79 |degrees~ C, the polycarbonate plate would deform as if its material stiffness were 40% of the room temperature modulus. Simply increasing the thickness of the plate with no ribs to 3.5 mm would provide a design that would meet the deflection requirements. The penalty would a be a 40% increase in material usage and an additional 8 sec in the cycle time. Choosing a material with more temperature resistance or initial stiffness is an option. On the other hand, a PC/ABS blend would have a reduced modulus of 10% of the room-temperature value, making it nearly impossible to consider that material.
Outdated measures of material performance such as HDT and notched Izod are inadequate for material selection or part design. Simple tools and techniques for predicting part performance (stiffness, strength/impact, and temperature resistance) integrated with processing concerns (flow length and cycle time) have been presented and demonstrated. Use of these tools is essential for optimum selection of materials and process conditions to meet demanding part performance requirements.
1. G.G. Trantina and D.A. Ysseldyke, SPE ANTEC Tech. Papers, 35, 635 (1989).
2. E.H. Nielsen, J.R. Dixon, and M.K. Simmons, ASME Computers in Engineering Conference, Chicago (July 1986).
3. K.C. Sherman, R.J. Bankert, and R.P. Nimmer, SPE ANTEC Tech. Papers, 35, 640 (1989).
4. H.G. deLorenzi and J.T. Woods, SPE ANTEC Tech. Papers, 39, 1411 (1993).
5. M.P. Sepe, SPE ANTEC Tech. Papers, 37, 2257 (1991).
6. D.O. Kazmer, SPE Injection Molding Division RETEC Proceedings, Boston (1990), p. K-1.
|Printer friendly Cite/link Email Feedback|
|Title Annotation:||Design; engineering plastics with cost effective design; optimum material and process selection|
|Author:||Trantina, Gerald G.; Oehler, Peter R.; Minnichelli, Mark D.|
|Date:||Aug 1, 1993|
|Next Article:||Concurrent engineering in product design/development.|