Interfacial Adhesion in Multi-Stage Injection Molded Components

Interfacial Adhesion in Multi-Stage Injection Molded Components Arvind Ananthanarayanan, Hugh A Bruck and S K Gupta, Department of Mechanical Engineering University of Maryland College Park, MD 20742, USA Abstract Several industrial components require assembly of multiple components to produce a final part. This leads to very high production and labor costs for the product. To overcome this, the use of multi-stage injection molding to produce articulated and compliant joints has been proposed. It is claimed that this method will produce in mold assembled components which would translate to reduced part counts and low assembly costs. Two broad classes of components that need to be produced using the multi material injection molding technology to facilitate in mold assembly are: 1) Components with articulated joints (e.g. Spherical joints, revolute joints etc.) 2) Components with compliant joints. During the multi-stage molding process, when the second material is injected on top of an already molded material, the two materials tend to adhere to each other. To create articulated devices, we will need to limit the adhesion at the interfaces so that we can obtain free moving articulated devices. At the same time, to produce compliant joints, we have to ensure appropriate strength at the interface to achieve in-mold assembly. Therefore, adhesion between materials during multi stage molding is one of the prime issues that needs to be dealt with while addressing the broader issue of multi stage injection molding for in-mold assembly. In this research investigation, mechanical testing and microstructural examination of the interface between the first and second stage parts has been conducted. The variation of interfacial strength with processing conditions, in particular the injection temperature, has been quantified and the primary mechanisms that affect interfacial strength identified as polymer crosslinking and shrinkage stress using two different specimen configurations, lap-shear and cylindrical pull-out, in order to determine the appropriate processing conditions for a given part geometry and desired level of interfacial strength. It was also determined that a lap shear specimen with a redesigned notch could provide insight into interfacial strength due to adhesion from polymer crosslinking, while the cylindrical pullout specimen could provide insight into geometric effects of shrinkage stress on interfacial strength. 1. Introduction Injection molding has emerged as the largest volume plastic processing technique for the mass production of plastic parts with complex, three dimensional shapes. By this method parts of good mechanical properties, complicated shape, made of almost unworkable materials, usually without additional machining can be manufactured. Several industrial components require relative movement between parts in order to fulfill a certain function. Traditionally such components were individually manufactured and subsequently assembled together in

order to fulfill their working requirements. But the assembly operation is a highly time consuming and labor intensive process which causes the production costs and the production cycle time to rise considerably. With the advent of injection molding as a popular process to produce plastic parts, it has therefore become imperative to explore new technologies which would enable manufacturing of movable parts using injection molding. Figure 1 shows an example of a rotor structure which was injection molded. The universal joint was assembled in mold using the overmolding operation. Figure 1 Rotor structure having two universal joints injection molded The rotor shown is molded in three molding stages to produce the two universal joints. The material used for all parts of the rotor is polyethylene. Therefore, in order to ensure relative motion between the various parts and hence the smooth functioning of the universal joints, it is imperative to ensure that there is no adhesion between the different parts of the rotor after the molding operations. Previous studies report adhesion between multiple materials and the conditions which would facilitate adhesion between components. Haberstroh et al have characterized the adhesion between Liquid silicone rubber (LSR) and thermoplasts following a two component injection molding operation. They have conducted several tests to come up with process parameters which would facilitate adhesion between the components [1]. Li et al have reported the characterization of interfacial failure in injection molded blends [2]. This study aims at establishing the various factors affecting interfacial adhesion between the polyethylene parts following the overmolding operation, and thereby suggesting ways to overcome this adhesion at the meso/macro scale. This study represents the first attempt to characterize interfacial strength in joints engineered through multistage injection molding where the objective is to minimize adhesion between the injection molded parts. These joints are manufactured using a new set of algorithms and analyses that have already been employed for joints fabricated from castable polymers [3-9]. 2. Qualitative Characterization of Interfacial Strength for Multi-Stage Injection Molded Components In order to better understand the processing conditions responsible for interfacial strength between molded low density polyethylene (LDPE) (T m =110 o C) components, a component with an articulating joint was designed. This part was used to conduct some preliminary qualitative testing. This part which contains two constrained revolute joints is shown in Figure 2. Figure 2 Rotary U-joint used for the pilot study

Step I Mold Stage 1 Step II Remove core Pour Stage 2 Step III Insert into stage 2 Step IV Step V Final Product Figure 3 Steps for molding Rotary U-joint In order to make the Rotary U-joint, the process steps involved are as follows: 1) The U-Rectangular hole part is molded in the first stage. A side core is used in the mold in order to make the hole in the part. 2) The core is removed from the part 3) The U-Rectangular hole part is inserted into the second stage mold. This part now acts as a mold for the second injection molding shot. 4) The rectangular pin is injection molded using a combination of the first stage part and a mold cavity as the mold. 5) The product is ejected from the mold. Figure 3 illustrates the process steps used for molding the Rotary U-joint using the overmolding process as described above. The melting point of the low density polyethylene used for the pilot study was 115 C. The Rotary U-joint was molded for different temperature ranges and for two different dimensions. The molding temperatures were varied from 115 C to about 140 C. The first stage and the second stage part were also molded at different temperatures in order to clearly establish the effect of temperature on the various molding stages. The principle observations that came out of this study are summarized as follows: a) The molding temperatures for the parts play an important role in the adhesion between the two stage parts b) The molding temperature for the second stage part plays a more pivotal role in determining the adhesion between the parts. The effect of the injection molding temperature of the first stage part is much lower in comparison. The observations from this study are tabulated in Table 1. S. No. First Stage Temperature ( C) Table 1: Variation of adhesion with temperature Second Stage Temperature ( C) Flange Thickness (inches) Pin Diameter (inches) Remarks 1 140 140.1.125 Complete adhesion. No movement possible 2 130 130.1.125 -do- 3 120 120.1.125 Little adhesion in the beginning. But when released, movement was achieved 4 115 115.1.125 No adhesion. Free movement. 5 120 140, 130.1.125 Complete adhesion. No movement

possible 6 115 140, 130.1.125 -do- 7 140,130 115.1.125 No adhesion. Free movement 8 140 140.25.25 Complete adhesion. No movement possible 9 130 130.25.25 -do- 10 120 120.25.25 Little adhesion in the beginning. But when released movement is achieved. 11 115 115.25.25 Erratic. Sometimes full adhesion. Lot of pores observed in first stage. 12 140, 130 115.25.25 No adhesion. Free movement. The initial study showed some very interesting results. But since the study was a predominantly qualitative study, it became exceedingly important to conduct a more quantitative study which would establish a clearer relationship between the injection molding temperatures and the adhesion between the two stage parts based on the adhesion forces involved. For this reason the lap-shear and the cylindrical specimen were studied for a more quantitative backing to establish the effect of temperature on adhesion. The results of this study are reported in the next section. 3. Quantitative Characterization of Adhesion Based on the findings of the pilot study, it was realized that the interfacial adhesion forces between the first stage and the second stage part need to be modeled on the basis of the second stage injection molding temperature. Therefore, additional quantitative characterization could be undertaken of the interfacial through standard mechanical testing. The quantitative interfacial adhesion study could be conducted by loading the specimen in tension, as well as shear. However, it was determined that the quantitative characterization is better carried out in shear loading since this reflected the loading condition that would be most indicative of the mechanical performance for revolute joints, such as the one seen in Figure 2. Hence a 0.2 thick and 0.4 wide specimen was modeled as a two stage injection molding product in a lap-shear geometry conforming to the ASTM D5868 standard, as shown in Figure 4 [10]. It is important to note that this standard was designed for testing the shear strength of adhesives that would be much lower than the strength of the parts that they joined, and that the types of joints being investigated in this study have unique interfacial characteristics that are a result of their processing. Therefore, the ASTM standard was also being investigated as a means for its validity for characterizing these types of joints. 1.2 0.75 0.05 Figure 4 Lap-shear component (All dimensions in inches) For the lap-shear component shown the injection molding steps followed were as follows: 1) The part shown in yellow was injection molded in stage 1. 2) This part was inserted into the stage 2 mold which, together with the mold cavity, acted as the mold for the stage 2 part. 3) The part shown in red was injection molded in stage 2. This component was molded at different temperatures and a shear failure test was conducted on these components in order to determine the adhesion forces between the two stage parts of polyethylene injection

molded together to produce the component shown above. This experiment was repeated for several temperatures between 120 C and 150 C. Three specimen components were prepared for each temperature and the shear loading test was conducted on each of these components. The failure loads were averaged over the three components. The results of this experiment are illustrated in Table 2. Table 2: Failure loads on shear for lap-shear component Temperature( C) Average Failure Load (lbs) 120 0.1 125 14.6 130 12.2 140 24.4 145 24.4 150 29.3 As can be seen from the graph in Figure 5, the adhesion strength or the corresponding interfacial failure shear stress roughly increases with temperature in a linear manner. But there are some variations in this trend at some temperatures. For example, at 130 C and at 145 C, the average failure load decreases. This irregularity can be attributed to some of the following reasons: 1) The loading may not in pure shear, and the bending stress may vary due to alignment of the specimen. 2) Molding flash causes some lateral adhesion which leads to a non pure shear loading. Also the parting line causes some molding flash which contributes to the adhesion between the first stage and the second stage part. 3) Also at higher temperatures the failure load is not indicative of the interfacial strength because the failure doesn t lie along the interface, as illustrated in Figure 6. Average Failure Load (lbs) 50 45 40 35 30 Failure at notch 25 20 Failure along interface 15 10 5 y = 0.8529x - 97.636 0 100 120 140 160 Injection Temperature (C) Figure 5 Variation of load with temperature for lap-shear component

(a) (b) Figure 6 (a) Failure along shear zone at t=130 C (b) Failure along stress concentration zone at t=150 C Although the above results show a reasonably clear trend regarding the relationship between adhesion and second stage injection molding temperature, a conclusive statement and quantitative prediction regarding the adhesion temperatures for different geometries can not be made due to the change in the failure mode that was observed. Hence it was felt that a new component design was required which would attempt to eliminate some of the problems caused by the lap-shear component and also help build a more confident estimate of the adhesion forces between the first stage and second stage component for any given geometry. Some of the requirements for the new component were: 1) Failure due solely to interfacial shear 2) Potential failure due to stress concentration on the components should be minimized 3) Variations in failure strength induced due to flash and parting line should be minimized For this reason pull-out specimen with a cylindrical hole was designed for further experimental tests. This specimen is shown in Figure 7. Figure 8 shows the processing steps for molding this component. 0.4 1.28 + 0.25 Diameter 0.28 Depth 1 Figure 7 Cylindrical hole specimen

Step 3: Pull-out core Step 1: Insert core into stage 1 mold Step 2: Inject stage 1 Final Part Step 5: Inject stage 2 Step 4: Insert stage 1 part into mold Figure 8 Processing steps for injection molding of pull-out component Some of the advantages of using the pull-out component for further adhesion tests were foreseen as follows: 1) A better estimate of pure shear adhesion could be got because of radial adhesion along the radial direction in the cylindrical specimen 2) Parting line problems can be minimized because of the presence of the side core in the molding operation because of which the hole can be produced without major errors. Similar tests as were conducted for the lap-shear component, were conducted for the pull-out component, where the surface of the hole in the solid piece from the first stage that was normal to the injection direction was coated with an anti-stiction coating to allow only the lateral surfaces of the hold to carry load in a shear state. Three specimens were tested for adhesion for each temperature between 120 and 155 C. The results of these tests are shown in Table 3. Figure 9 shows the relationship between adhesion forces and second stage injection molding temperatures for the pull-out component, which follows a linear trend that is similar to the lap-shear configuration. Temperature C Avg Max load(lbs) 120 4.9 125 14.6 130 23.2 135 22.8 140 15.46 145 29.3 150 29.3 155 41.5 Table 3: Failure loads on shear for cylindrical hole component

Average Failure Load (lbs) 50 40 30 y = 0.8111x - 88.893 20 10 0 100 120 140 160 Injection Temperature (C) Figure 9 Variation of load with temperature for cylindrical hole component As can be seen from the trend, the variation of injection molding temperature with the adhesion appears to be in reasonable agreement with the results for the lap-shear component. Some irregularity can still be seen between the temperatures of 130 C and 140 C. Some of the reasons that can be attributed to this irregularity are: 1) The first stage part tends to shrink after the injection molding which leads to considerable flash in the second stage part. This flash tends to cause additional adhesion between the parts which tend to distort the results obtained. 2) The parts tend to adhere to the polyurethane molds after the injection molding operation. The adhesion force between the two parts tends to reduce due to the force applied to the component during ejection. 3) Due to shrinkage of the first stage component a bulge in the second stage component is noticed. This causes geometrical interlocking between the first stage and second stage component which contributes to the adhesion forces. Figure 10. Pull-out component after failure Figure 10 shows the failure for the pull-out component which clearly illustrates the abovementioned points. The statistical variance is observed to be most between the temperatures of 130 C and 140 C. For this reason some irregularity is observed in the trend for these temperatures. 4. Discussion The two sets of experiments conducted above give a clear indication that the adhesion forces in injection molding increases with the injection molding temperature. The least moldable temperature, which is close to the melting

point of the polymer (around 120 C for polyethylene), therefore, gives best results for the manufacturing of articulated joints. Figure 11 shows the results of the two sets of experiments superimposed together in order to compare the two testing configurations. In order to understand geometric effects on the resistance to failure, the results are normalized by the contact area of the interface to convert the load to an averaged failure stress over the interfacial area. It is interesting to note that the loads for the cylindrical specimens were approximately 1/3 that of the lap-shear specimens. This would tend to indicate some kind of geometric effect on the adhesion mechanisms. The lack of a transition in the failure from shear to normal in the pull-out specimen, which would be expected for a single shot specimen, would also indicate that the strength of the cylindrical interface is not as great as the lap-shear at the higher injection temperatures where the failure transitioned to the notch from the lapshear interface. Averaged Shear Failure Stress (lbs/sq. in.) 400 350 300 250 200 150 100 50 0 Cylindrical pull-out Lap-shear y = 3.6883x - 404.22 y = 10.661x - 1220.4 100 120 140 160 180 Injection Temperature (C) Figure 11. Comparison of results for the cylindrical hole and the lap-shear specimen There are three mechanisms that are responsible for the interfacial strength in these specimens: a) Adhesion from polymer crosslinking. b) Plastic welding at interface. c) Geometrical interlocking. Polymer crosslinking requires chemical activity between the surfaces, which is usually obtained in the melt state. This is possible at higher injection temperatures, where the injected melt can possess enough energy to locally melt the solid piece in the first stage. If the melting is not sufficient to activate the surfaces, it is still possible for the surfaces to substantially deform near the interface and have the two surfaces intertwine to create plastic welding. Finally, if there is sufficient roughness or curvature at the interface, it is possible for the two solid piece in the second stage to interlock with the solid piece in the first stage. Each of can be isolated, since mechanical interlocking can only exist at lower melt temperatures where there is not sufficient energy to activate or melt the solid piece in the first stage. Thus, a straight interface with a smooth finish, as is produced after the first stage, will eliminate this mechanisms. When there is sufficient energy to melt the solid piece in the first stage, distortions of greater than 1 micron in the straight interface should be visible under an optical microscope. Distortions below this level induced by roughening the surface of the solid piece in the first stage with sandpaper have typically shown to have negligible effect on the interfacial strength. Therefore, when there is interfacial strength and no distortions appear at the interface, then the effects can be attributed to polymer crosslinking.

In order to determine if there is plastic welding at the interface, a study of the interface was conducted under the microscope for the lap-shear component. Different colors were used to distinguish between the first stage and the second stage part. The interface was subsequently polished. For magnifications as high as 10X, no weld pool was observed at the interface. This is illustrated in Figure 12. For this reason, plastic welding at the interface ould be ruled out. Geometrical interlocking could not have been a reason for the adhesion in the lap-shear specimens, since the interface between the first stage and the second stage parts is free. However, geometrical interlocking could play a prominent role in the adhesion between the first stage and the second stage of the cylindrical hole pull-out specimens. However, the lower strength of the cylindrical interfaces would tend to rule this out. Instead, there is a substantial shrinkage (~1-2%) of these polymers as they cool. This shrinkage would be radial in the cylindrical specimens and lateral in the lap-shear. Thus, stress due to shrinkage would more likely contribute to debonding in the cylindrical specimen and not in the lap-shear. However, using an elliptical failure criterion with the normal failure stress approximately 3 times that of the shear failure stress, the radial shrinkage stress would have to be about 94% of the tensile failure stress to reduce the shear stress by 1/3. This is quite a large value to obtain from 1% shrinkage, and would greatly comprise the strength of the interface. Nevertheless, it is still possible for substantial stress to accumulate and influence the interfacial strength. Hence, it can be said that a combination of shrinkage stress due to geometric effects and polymer crosslinking could be a possible explanation for the adhesion between parts molded in subsequent stages using injection molding operation, and that the lap-shear configuration is more appropriate for studying the crosslinking effects with a redesign of the notch, while the pull-out configuration could provide insight into shrinkage effects. Is also possible that geometric interlocking could be added to the interfaces to quantify their contribution to interfacial strength. Such a study has been previously conducted on cast multi-stage molded interfaces, where the failure strains associated with a transition from interfacial fracture controlled by polymer crosslinking to ligament failure controlled by the geometrical interlocking were characterized using Digital Image Correlation (DIC) [3]. Figure 12. Interface between first stage and the second stage parts studied under a microscope with 10X magnification (Pink = First stage section, White = Second stage section). 5. Conclusions and Future Work 500 m A new process for designing joints with controlled interfacial adhesion through multi-stage injection molding has been developed. The conditions for controlling the interfacial adhesion was first qualitatively characterized for a model revolute joint using polyethylene as a model material, and determined to be related to the injection temperature for the second shot. A quantitative study was then conducted using an ASTM standard lap-shear geometry. These tests indicated a logarithmic trend in adhesive strength from 120 to 150 o C. However, the design of the specimen was for characterizing the shear strength of adhesives that were much lower than the parts they joined. Consequently, the trend reflected a change in failure from the interface to a stress concentration that was located near a notch in one of the base components at the highest failure load. Therefore, a second specimen geometry was designed based on a pull-out test. These second tests also produced a linear trend that was reflective of a more consistent failure mode over the same temperature range. However, the interfacial strength was reduced by 1/3, and was attributed primarily to radial shrinkage stresses. Thus, it was determined that a lap shear specimen with a redesigned notch could provide insight into interfacial strength due to adhesion from polymer crosslinking, while the cylindrical pull-out specimen could provide insight into geometric effects of shrinkage stress on interfacial strength.

6. Acknowledgements This research has been supported in part by NSF grants DMI0093142 and DMI0457058, the Army Research Office through MAV MURI Program (Grant No. ARMY-W911NF0410176), and by the National Institute of Aerospace, Hampton, VA. Opinions expressed in this paper are those of the authors and do not necessarily reflect opinions of the sponsors. 7. References [1] E. Haberstroh and C. Ronnewinkel. LSR thermoplastic combination parts in two-component injection molding, Journal of polymer engineering, 21, 303-318, 2001. [2] Zhong-Ming Li, Zhi-Qi Qian, Ming-Bo Yang, Wei Yang, Bang-Hu Xie, and Rui Huang. Anisotropic microstructure-impact fracture behavior relationship of polycarbonate/polyethylene blends injection-molded at different temperatures, Polymer, 1-12, 2005. [3] H.A. Bruck, G. Fowler, S.K. Gupta and T.M. Valentine. Using Geometric Complexity to Enhance the Interfacial Strength of Heterogeneous Structures Fabricated in a Multi-stage, Multi-piece Molding Process, Experimental mechanics, 44, 261-271, 2004. [4] R.M. Gouker, S.K. Gupta, H.A. Bruck, and T. Holzschuh. Manufacturing of multi-material compliant mechanisms using multi-material molding. International Journal of Advanced Manufacturing Technology, 28, 1-27, 2006. [5] A. Priyadarshi and S.K. Gupta. Geometric algorithms for automated design of multi-piece permanent molds. Computer Aided Design, 36(3):241 260, 2004. [6] X. Li and S.K. Gupta. Manufacturability analysis of multi-material objects molded by rotary platen multi-shot molding process. In ASME International Mechanical Engineering Congress and Exposition, Washington, DC, 2003. [7] X. Li and S.K. Gupta. Geometric algorithms for automated design of rotary-platen multi-shot molds. Computer Aided Design, 36, 1171 1187, 2004. [8] M. Kumar and S.K. Gupta. Automated design of multi-stage molds for manufacturing multi-material objects. ASME Journal of Mechanical Design, 124, 399--407, 2002. [9] S. K. Gupta and G. Fowler. A step towards integrated product/process development of molded multimaterial structures. In Tools And Methods Of Competitive Engineering, Lausanne, Switzerland, April 2004. [10] ASTM standard D5868. Standard test method for lap-shear adhesion for fiber reinforced plastic (FRP) bonding, 2006.