Sequential injection molding of thermoplastic polymers. Analysis of processing parameters for optimal bonding conditions

Sequential Inject ion Molding of Thermoplastic Polymers. Analysis of Processing Parameters

for Optimal Bonding Conditions

A. R. CARELLA’*, C. ALONSOl, J. C . MERIN0’g2, andJ. M. PASTOR’s2

Center for Automoti~~e Research and Development (CIDAW Technological Park of BoeciZlo 47151 Valladolid, Spain

2Dept. Fisica de la Materia Condensada, E.T.S.I.I. University of VaUadolid 4701 1 Valladolid Spain

A systematic approach was used to study the effect of the process variables that control bonding of the injected melt to the previously injected parts, in sequential injection molding of thermoplastic polymers. Three polymer pairs were sequentially injected in a mold with the same geometry, in a range of mold and injected melt temperatures, and packing pressure conditions. Standard Peel Tests were conducted on injected samples to measure the resulting bonding strength. The analysis of the experimental results allows the quantification of the relative importance of the processing parameters involved. A direct correlation is found between the calculated interface temperature and the packing pressure needed for bonding. A generalized procedure is proposed to establish the processing conditions, which allow polymer- polymer bonding to be optimized.

INTRODUCTION to take care of the user’s comfort or to admit larger

equential injection molding of thermoplastic poly- S mers has gained considerable importance in re- cent years (1). There is some literature on this subject (2) for vitreous polymers, but the bonding mechan- isms, as well as the testing geometry used and the thermal histories, are different.

In a single mold with two or more inlets and moving parts that allow sequential entrance of melts into different sections of a main cavity, two or more types of thermoplastics are sequentially injected. The final product is a single piece made of polymers with different mechanical properties, which are permanently bonded together.

A large fraction of the articles manufactured by sequential injection molding is used to provide comfort to users. Most of the applications involve covering a rigid polymer with a soft elastomer, for car parts and household appliances. The rigid polymers are used for struc- tural purposes, and the softer materials are supposed

deformations without noticeable damage. Many &- ples of these applications can be seen as knobs, steer- ing wheels, tools, and household appliance handles and grips, and so on. Older production systems con- sisted of injection molding all components separately, and later assembling and gluing or welding them to make the final article; sequential injection saves production costs and increases productivity.

Complex articles can be obtained with the sequential injection molding process. A fundamental requirement is that all the sections that form these articles are well bonded together. Correct bonding is necessary to sat- isfy mechanical and aesthetical requirements.

During sequential injection molding, a thermoplastic melt is injected into a cavity closed by walls that may be part of the mold or part of the final article being molded. Cooling of the melt at the mold walls is a well-known process. The injected melt must bond to the previously injected parts (made of different thermoplastics), and then cool to solidify.

Throughout this work, a systematic approach was used to study the effect of the process variables that

- control bond& of the injected melt to the previously injected parts and subsequent cooling process. Com- puter simulations were performed with a commercial

*To whom mmpondence should be addressed. present address: Dpto. hgenhia de Materiales. Facultad de Ingenieria. Universidad Nadonal de Mar del !‘lab Juan B. Just0 4302 .(7600) Mar del Plata. Republica Argentina.

2172 POLYMER ENGINEERING AND SCIENCE, NOVEMBER 2002, Vol. 42, No. 11

Sequential Injection Molding

program (MoldFlow) in order to predict temperature and pressure profiles at the melt-solid polymer interface for an adequately simple geometry. Three polymer pairs were sequentially injected in a mold with the same geometry, in a range of processing conditions suggested by the simulation results. Peel tests were conducted on injected samples to measure the resulting bonding strength. The analysis of the experimental results helps iden* the relative importance of the processing parameters involved. A generalized procedure is proposed to establish the processing conditions that allow polymer-polymer bonding to be optimized.

EXPERIlUENTAL

Experiments were conducted on three pairs of thermoplastic polymers. For the sake of simplicity, only work done with one of these polymer pairs will be fully described.

Two types of Sarlink TPO were used for these experimental studies. Sarlink is a trade name for a family of polyolefin thermoplastic elastomers (TPO) that are obtained by dynamic vulcanization. This material consists of a continuous polypropylene (PP) matrix and a dispersed crosslinked rubber (EPDM) phase. Several grades are available in the market, covering a wide range of properties like hardness, oil resistance, compression-set values and others. The grades used were Sarlink 4149 and Sarlink 4165B, which will be referred to as S1 and S2. These - 0 s show different hardness and oil resistance values. The number 65 stands for 65 Shore A hardness and the number 49 stands for 49 Shore B hardness. The letter B identifies the black color. Typical properties for both grades as given by the manufacturer are listed in Table 1.

Part of the data needed for injection computer simulations such as thermal conductivity (k) data for the temperature range used (Table 1). pressure-volume- temperature (PVT) curves covering the range between LO5 and 1.6 X lo8 Pa, and between 298 and 498 K, and viscosity data covering the range between 100 and 10,000 sec-l, and between 453 and 493 K were provided by the manufacturer.

Melting temperatures and specific enthalpy were measured by DSC (Mettler-Toledo 821/400). A first scan was run h m ambient temperature to 523 K, to erase previous thermal and stresses history; the sample

was then cooled to 173 K, at about 10 K/min; the useful data was then recorded in a second heating from 173 K to 523 K, done at 10 K/min. Specific heat (Cp) values for several temperatures were calculated from DSC data.

Sample Preparation

The injection molding machines used for the sequential injection molding process consist of two or more injection barrels filling a single mold with mov- able parts. Since this type of equipment was not available, samples suitable for this work were prepared by using a single-barrel injection molding machine and a specially designed mold.

The injection molding machine used is a Margarit JSW 110. Its most important features are: screw di- ameter: 0.045 m. Maximum injection pressure: 1.5 X lo8 Pa. Maximum injection volume: 2.23 X m3. Maximum injection speed: 1.43 X m3 X sec-'. Maximum clamping force: lo6 N.

The samples made of two different materials were obtained by a two-step process, which involved first injecting white S1 into a thin rectangular mold cavity (0.15 m X 0.047 m X 0.002 m measured along the x, y, and z axes). The injected pieces were then placed in a double thickness (0.15 m X 0.047 m X 0.004 m) mold cavity, and allowed enough time to equilibrate at the mold temperature (about 30 min). Carbon black- colored S2 was then injected into the remaining cavity. Both materials are injected at the lefbside end of the mold. A short non-bonded interface length (necessary to attach the testing machine clamps for peel tests) was produced at the right-side end of the molded samples by adhering a (0.040-m-long) thin aluminum tape to the S1 part before the second injection step.

Table 2 lists molding conditions for all specimens tested. The melt injection temperatures were chosen low enough to prevent any thermal degradation from occurring. The packing pressure values used were always restricted to the range that ensured the post- molding dimensional stability of the composite injected parts.

Peel Temt.8

Bonding strength of the S 1-S2 interfaces was measured, as suggested in ASTM D 903 - 93 (3). by a 180" peel test at a peel rate of 0.025 m/min using an In- stron Universal Testing Machine Model 5500 R 6025.

Table 1. Phvsical Properties of the Materials Used.

s1 2.47 0.08 b 0.92 433 a 413-433 a

s2 2.39 a 0.057 b 0.93 b 433 a 413-433 a

PP 2.14 0.16 0.91 433 EPDM + Oil 3.4 0.22 0.95

a- Calculated from DSC curves. b- Provided by manufacturer. c- Obtained from literature.

POLYMER ENGINEERING AND SCIENCE, NOVEMBER 2002, Vol. 42, No. 11 2173

A. R. CareUa, C. Abnso, J. C.

Table 2. Molding Conditions for All Specimens Tested.

Melt Injection Mold Temperature Packing pressure Temperature (K) (K) (MPa)

463 318 20 30

338 20 30

473 31 8 20 30

338 20 30

478 31 8 20 30

328 30

333 30

338 20 30

483 31 8 20 30

323 20 30

333 20 30

338 20 30

488 31 8 20 30

323 30

333 30

338 20 30

493 31 8 20 30

338 20 30

Since the non-bonded interface length necessary to carry on the 180" peel test is 95 mm (larger than the non-bonded interface length obtained using the aluminum tape), an extra length was obtained separating both layers in a 90" peel setup at 0.010 m/min. There- fore, the total length peeled on the 180" peeling test was always 0.055 m.

The S2 layer was always backed up with a 0.1-mm- thick PET film, adhered with a cyanoacrylate glue ( h c - tite 4062, over a S2 surface previously primed with Loctite 770) to prevent elongation of the softer layer during the peel test. b e n t i n g elongation of the softer layer during tests justifies the calculation of the peeled length as half the crosshead displacement.

The load was recorded as a function of crosshead displacement while both layers were peeled apart (3). The peeling force was calculated as P / w , where P is

Merino, and J. M. Pastor

the applied load and w is the width of the specimen. Peeling force defined in this way corresponds to the energy required per unit area of interface peeled. After peeling, all specimens were measured in order to determine the position of the point at which the failure mode changes from adhesive to cohesive. Adhesive failure occurs when the interface is weaker than any of the layers; cohesive failure occurs when the interface is stronger than the weakest layer. Adhesive failure is easily identified for these TPOs because the peeled surface of the stronger (white) layer is obtained clean of any traces of the weaker (black) layer. Cohe- sive failure is defined for these peeling tests as the peeling mode that leaves at least 50% of the stronger layer surface covered with small pieces of the peeled weaker layer.

Computer Simulations

Commercial simulation software (MoldFlow) was used to calculate temperature and pressure profiles in the 52 liquid phase, for the filling and cooling stages. Thermodyrmmc and transport data used are included in Table 1.

DISCUSSIOIIT OF PEEL TEST RESULTS mure 1 shows typical raw results obtained from a

peeling test. The recorded pew force is plotted versus the crosshead displacement (twice the peeled length). The trace shows oscillations that are characteristic of peeling tests of elastic materials, as these do not prop- agate peeling at a constant rate. For this reason, an average peeling force value is used for calculation purposes. For displacement values below 0.060 m, the average peeling force measured for this specimen is 1950 N/m; for displacement values above 0.065 m, the average peeling force is 3500 N/m; this value corresponds to the tearing energy per unit surface for S2 for this testing geometry, and should be the same for all tested specimens, within experimental error.

For the previously mentioned experiment, visual in- spection after peeling reveals a clean S1 surface for peeled lengths below 0.030 m (half the crosshead displacement for 1950 N/m average peeling force). The continuation of the peeled surface shows a short transition zone (about 0.002 m long) and then the zone where the failure is cohesive, covered by a thin layer of tom black S2; this zone corresponds to the average peeling force of 3500 N/m.

The 180" peeling test geometry is able to measure peeling forces up to the tearing energy per unit surface for the weaker material. Any stronger bonding at the interface will not be detected by this testing geometry. For this reason, we have defined bonding at the interface as belonging to one of two categories-either insu-t bonding, which corresponds to peeling forces lower than the tearing energy per unit surface for the weaker material, or sujjWent bonding, which corresponds to peeling forces equal to the tearing energy per unit surface for the weaker material.



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The above-mentioned experimental observations in- dicate that the bonding process of the S2 on the solid S1 is conducted under conditions that change from the left-side end of the mold to the right-side end. The melt fills the mold completely in about 0.4 second, starting from the left-side end. Moldnow simulations show that the melt reaches the solid polymer surface at temperatures that do not differ by more than 3 K, from one mold end to the other. This is mainly due to the viscous heating, which compensates for heat losses. Simulations also show that the packing pressure decreases markedly along the melt path, giving a linear pressure profile from the left-side end of the mold toward the right-side end.

Another useful observation is that bonding for specimens injected at constant mold and melt temperatures can be improved by increasing the packing pressure. As a general rule, increasing packing pressure increases the peeling force for the zone where the failure is adhesive, and also shortens its length. The effect of increasing packing pressure from 20 to 30 Mpa can be observed in FQ. 2. The average peeling force for the zone where the failure is adhesive in sample W is about 2 125 N/m and the length of the adhesive failure zone is about 0.030 m. The average peeling force for the zone where the failure is adhesive in sample V (packed at the higher pressure) is about 3200 N/m, and the length of the adhesive failure zone is about 0.022 m. For these mold and melt temperatures, increasing packing pressure by 50% increases the average peeling force by 40%, and shortens the adhesive failure zone by 25%. These effects of changing packing pressure also depend on mold and melt temperatures. For samples where the zone of adhesive failure is long

enough, a smooth change can be detected for the peeling force along the adhesive failure zone, which is assumed to be due to the linear increase of the applied packing pressure. Flgure 3 shows an example of this behavior.

m e 4 shows peeling force for the adhesive failure zone, versus melt injection temperature for several mold temperatures and two packing pressure values. The effects of mold temperature and packing pressure are more evident in the range of temperatures between 473 K and 490 K. For constant melt injection temperature, the effects of increasing mold temperature and increasing packing pressure are shown as increasing peeling force for the adhesive failure zone. For a melt injection temperature of 483 K, raising the mold temperature from 318 K to 338 K increases the peeling force by a factor of 7. For the same melt injection temperature, increasing the packing pressure by 50% also increases the peeling force by factors that depend markedly on mold temperatures. For melt injection temperatures lower than 463 K, or higher than 493 K, the effects of mold temperature and packing pressure are considerably reduced: this effect will be explained later. It is expected that-for the lower melt injection temperatures used-further increases in mold temperature will increase the peeling force and reduce the slopes of the curves shown in Flg. 4. Also-for the higher melt injection temperatures u s e d - w e r re- ductions in mold temperature will reduce the peeling force and increase the slopes of the curves shown in Flg. 4.

Flgure 5 shows measured lengths of cohesive failure zones, versus melt injection temperature for several mold temperatures. These values were obtained by


A. R. Carella, C. Alonso, J. C. Merino, and J. M. Pastor

For both specimens: Melt Injection Temperature = 483K

0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14

Crosshead displacement (m)

(a)

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For both specimens: Melt Injection Temperature = 483K Mold Temperature = 338K Packing Pressure = 30 MPa

I - Specimen v1 Specimen V2

-Linear fit of adhesive failure zone 0 I I I I I I I

0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14

Crosshead displacement (m) (b)

Fig. 2. Peeling force versus crosshead displacement plotted for applied pachxtg pressures of 20 MPa (A) and 30 MPa @).

measuring the tested specimens that provided the data used for Q. 4. Since the total interface length peeled in the 180" peel test mode is always 0.055 m, an increase in the length of the cohesive failure zone corresponds to a reduction in the length of the adhesive failure zone. Again, the effects of mold temperature and packing pressure are more evident in the range of temperatures between 473 K and 498 K. It is observed that a rise in mold temperature and packing

pressure produces an increase in the lengths of cohesive failure zones. In the same way, they influence the peeling force results.

After these general observations about the effects of the injection variables on the bonding quality obtained for different parts of the injected composite samples, one point must be stressed: the objective of this work is to develop a generalized optimization procedure that can be applied to any thermoplastic pair to



For both specimens: Melt Injection Temperature = 483K Mold Temperature = 323K

2 2500-

a 2000-

0.00 0.02 0.04 0.06 0.08 0.10 0.12 0.14

Crosshead displacement (m) Flg. 3. Raw peeling test data showing variations of peeling force along the adhesive failure zone.

-*-- Pp 30MPa MT 323K 43- Pp 30MPa MT 328K

450 455 460 465 470 475 480 485 490 495 500

Melt Injection Temperature (K) Fig. 4. Peeling force uersus melt injection temperature curves for dipwent mold temperatures and packing pressures.

be sequentially injected. The criterion used here is to minimize the required mold temperature, melt injection temperature, and packing pressure, while obtaining cohesive failure for the whole interface. Therefore, the lower limits for the above-mentioned injection variables are taken as those that generate the transition mne between adhesive and cohesive failures.

Ngure 6 contains a general view of the results obtained for the ranges of packing pressure, mold temperature, and melt injection temperature used for this

work. The abscissa shows melt injection temperatures for all the samples. The ordinate shows the calculated local packing pressure corresponding to the length (measured along the melt path inside the mold) at which the transition zone is produced for each sample. Lines join results obtained for samples made at the same mold temperature. Results obtained for two different overall packing pressures coincide in the same joining lines. Note that overall packing pressure is different from local packing pressure, which varies


A. R. CareUa, C. Alonso, J. C. Merino, and J. M. Pastor

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A e e l i n g start -0- Pp 30MPa. MT 318K -0- Pp 30MPa, MT 323K -A- Pp 30MPa, MT 328K -0- Pp 30MPa, MT 333K

-x- Pp 30MPa, MT 338K + Pp 30MPa, MT 335.5K

X . 0.000~ - 1 . I . I . , . I . , . , . , , I .

450 455 460 465 470 475 480 485 490 495 500

Melt Injection Temperature (K) Q. 5. c~hesiue f i u r e kngth versus melt injection temperature curves for di3ment mold temperatures at 30 MPa of *plied packing presswe.

along the mold length, and the coincidence is due to the coincidence of local packing pressure.

Flgure 6 can be used to predict the results for injecting S2 over S1. Any operation condition placed above or to the right side of the line corresponding to the mold temperature used will result in sufficient bonding. Processing at the minimum temperature and pressure conditions will reduce thermal degradation and energy requirements.

SELECTION OF OPTIMUM PROCESSING PARAMETERS

A general approach is proposed here. An equivalent series of inexpensive experiments must be performed in a mold similar to the one described here. The injections must cover the expected ranges for mold temperature, melt injection temperature, and packing pressure. Peeling experiments give information on the peeling forces and cohesive bonding lengths. Either

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450 455 460 465 470 475 480 485 490 495 500

Melt Injection Temperature (K) Rg. 6. Local packing pressure versus melt injedion temperature c u ~ e s for diserent mold temperatures.



POLYMER ENGINEERING AND SCIENCE, NOVEMBER 2002, Vol. 42, No. 11

data can be used for the proposed method, while obtaining both will give a safer, self-consisting procedure. These experiments give the needed experimental data for the pair of thermoplastic polymers to be sequentially injected. Packing pressures can be calculated with simulation software. A graph similar to Rg. 6 can then be constructed, and used to establish the lower limits for injection conditions that will give sufficient bonding. For other parts with different geometries, simulations can then be performed to determine the machine operation conditions needed for safe production.

Some guidelines for a primary selection of melt injection temperatures and mold temperatures can be learned from simple considerations. Cohesive bonding will be obtained via the semicrystalline polypropylene phase, and for this purpose, thin polypropylene layers next to the interface must be put in contact in the melt state. From literature data on diffusion rates for the molecular weights corresponding to commercial polypropylene (4, 5), we can be sure that no apprecia- ble transport of center of mass will occur across the interface within the processing times used for injection molding 16). On the other hand, some molecular interdigitation will take place above melting temperatures, and subsequent co-crystallization upon cooling will provide effective bonding. Therefore, the thickness of melted solid layer next to the interface need be only a few radii of gyration (less than 100 nm for a commercial polypropylene). For this to occur, the injected melt will have to provide enough enthalpy to melt a small solid layer thickness while remaining in the melt state. A quick estimation of interface temperature can be calculated from standard heat transfer equations for two bodies put in contact at different starting temperatures (7):

Ti * bl + 2-2 * 4 Ti = (1)

bl + 4 Here, TI and T2 stand for the melt injection temper-

ature and the mold temperature, and b, and b, corre- spond to the thermal effusivities of both materials (k1/2*p1/2*Cp1/2); Ti is the interface temperature obtained after contact.

Equation 1 is obtained under the assumptions of unidirectional heat transfer between two infinitely thick bodies separated by an infinitesimally thin sheet that results from the perfect mixture of equal amounts of each layer. The obtained results are applicable for short times ( t < 5 seconds) that are larger than the typical injection time. This method is always approxi- mate because Cp is assumed a constant, and crystallization or melting enthalpies are not considered. A more elaborate calculation scheme is being tested, and will be included in a future publication.

Preliminary calculations based on temperatures included in TabZe 2, for conditions where cohesive bonding was obtained, gave some surprising results, because the Ti values fell below the S1 melting temperature range. These calculations were done assuming that

the thermal efisivity values were similar for the melt and the solid; this procedure is equivalent to consider both as single-phase materials. As the polypropylene phase and the oil-filled cross-

linked EPDM have different thermal effusivity values, another approach can be used, assuming that the EPDM phase of the melt is put in contact with the solid polypropylene. For this situation, the interface temperatures fall well into the S 1 melting temperature range. This seems to be a more appropriate way to es- timate the maximum interface temperatures.

In Fig. 5 we can see that-at a constant packing pressure-the lengths of the cohesive bonding zones increase when the calculated interface temperature is increased. We can also see that the same cohesive bond length can be obtained for different combinations of melt and mold temperature, as long as the calculated interface temperature remains the same. Furthermore, the calculated interface temperatures for specimens that were injected with the same packing pressure and show equal lengths of cohesive failure zones fall within narrow ranges. Calculated interface temperatures for all injected specimens are included in Table 3.

In Rg. 7, interface temperatures (IT) (calculated fol- lowing the above-suggested procedure) have been plotted versus local packing pressures (UP). In this range of temperatures and pressures, an almost linear correlation is obtained and shown in Eq 2.

As mentioned before, for melt injection temperatures lower than 463 K or higher than 493 K, the effects of mold temperature and packing pressure are considerably reduced; a qualitative explanation can now be suggested for these effects. When the hot melt is placed in contact with the cooler solid polymer, temperature gradients for the solid and the liquid are generated by heat conduction (7)- At temperatures close to the melting point, the temperature gradient in the liquid will cause differential thermal contrac- tions, larger at the liquid surface and smaller in the bulk. These contraction processes may well lead to the formation of cavities at the original liquid-solid interface. Melt-solid heat transfer is considerably im- paired by cavities. Enough packing pressure may prevent cavity formation, keeping good heat transfer characteristics between melt and solid, and allowing the solid to receive enough heat to melt the necessary thin solid layer, as explained before. When the melt injection temperature is too low, there is no amount of packing pressure that can cause sufficient bonding, because there is not enough melted solid at the interface. When the melt injection temperature is high enough, the cavities cannot be formed, because enough solid layer thickness is melted, and interdigitation takes place before the original liquid layer temperature is lowered enough to start differential contraction.

2179

A. R. CareUa, C. Alonso, J. C. Merino, and J. M. Pastor

428 - 426-

424 - 422-

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5 416: Interface Temperature = 441.7 K - 1.73 (K/MPa)*Local Packing Pressure 3 414-

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Table 3. Calculated Interface Temperatures for All Mold and Melt Injection Temperatures.

I Kl 31 8 323 328 333 338 Mold temperature 4

Melt injection temperature (K)

4 Interface temperature [K]

463 404 406 408 410 41 2 468 407 409 41 1 41 3 41 5 473 41 0 41 2 414 41 6 41 8 478 41 3 41 5 41 7 41 9 421 483 416 41 8 420 422 424 488 419 421 423 425 427 493 422 424 426 428 430 498 425 427 429 431 433

It is important to notice that the operation temperatures (mold and melt) as well as the packing pressure values used were always restricted to a range that prevented degradation and ensured the post-molding dimensional stability of the composite injected parts. Because of the above explanation, no higher packing pressures were used, as they were assumed not to be useful.

CONCLUSIONS These results suggest a standard procedure for

choosing melt and mold temperatures that will give sufficient bonding. First, suitable combinations of mold and melt temperatures can be selected to obtain interface temperatures within or above the melting temperatures range of the solid layer. Then, a small set of experimental injections in a range of packing pressure values can be performed in a small mold. One-hundred-eighty degree (180") peeling tests will

provide the necessary data to construct a graph similar to Q. 7, which will provide the information needed to establish the optimum processing conditions. These results can be extrapolated to other mold geometries by simulation.

Interface temperature calculations based on Eq 1 also suggest that the heat transfer rate across the interface plays a very significant role in the whole bonding process. In this sense, it can be assumed that poor heat transfer between melt and solid may decrease or prevent the solid surface melting, and impair the liquid interdigitation, causing a poorer bonding.

This form of optimization for the injection processing conditions may play a very important role in improving the quality of the finished product. Melt temperatures may be reduced with appropiate combinations of mold temperature and packing pressure, to control warpage and/or degradation: lower mold temperatures may speed up molding times, etc.

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9.0 9.5 10.0 10.5 11.0 11.5 12.0 12.5 13.0 13.5 14.0 14.5 15.0

Local Packing Pressure (MPa) Fig. 7. Interfime temperature versus local packing pressure.

21 80 POLYMER ENGINEERING AND SCIENCE, NOVEMBER 2002, Vol. 42, No. 11


ACKNOWLEDGMENTS 3. Standard Test Method for Peel or Stripping Strength of Adhesive Bonds, ASTM D 903 - 93.

4. J. Karger-Kocsis, Polypropylene Structure, Blends and Composites, Chapman and Hall, London (1995).

5.T. Schuman, E. V. Stepanov, S. Nazarenko. G. Capaccio, A. Hiltner, and E. Baer, Macromolecules, 31, 4551 (1998).

6. N. 2. Qureshi, E. V. Stepanov, G. Capaccio, A. Hiltner, and E. Baer, MacromotecUtes, 34, 1358 (2001).

7. Agassant, Avenas, Sergent. and Carreau, Polgmer Process- ing. principles and Modeling, Hanser Publishers (1991).

Financial support from CICYT (1FD 1997-2 115- C02.02) and CYTED Cvn1.4) is acknowledged.

REFERENCES 1 . Modem Plastics, 71, 354 (1994). 2. D. Y. Huang and R. S. Chen, Polym. Eng. Sci., 39, 2159

( 1999).

POLYMER ENGINEERING AND SCIENCE, NOVEMBER 2002, Val. 42, No. 11 2181

Documents

Sequential injection molding of thermoplastic polymers. Analysis of processing parameters for optimal bonding conditions