volume 35 winter 2010 message from the extrusion division ... · extrusion problems with flow...

Winter 2010

Volume 35

Number 3

Inside this issue:

Extrusion

Wiki, ANTEC

2011

2

Aalysis of

Extrusion

Problems

with Flow

Simulation

3

Bulk Resin

Feed 9

ANTEC 2000

Best Paper

Flow Surging

in Single

Screw Ex-

truder

15

Extrusion

Hints

8,

12,

13,

14,

22

Extrusion

Division

Board

25

Message from the Extrusion Division Board of Directors Chair

Fellow SPE and Extrusion Division current and future members,

As the Great Recession slowly recedes into the realm of bad memories, business

has made a turn for the better. We can once again return to some semblance of

"business as usual". That said, the Extrusion Division board of directors has been

very busy to provide services and information to you, our all important mem-

bers. The Newsletter has returned with improved content and frequency, thanks

to the heroic efforts of Gary Oliver and Karen Xiao. The Extrusion Topcon will

be held in Charlotte, N.C. on November 10 and 11. A full and robust technical program is being

organized for the ANTEC 2011 in Boston. Contact Michelle Curenton for more informa-

tion. The Extrusion Division website and WIKI continue to add content and provide links to

high quality information and contact information. Extrusion Division sponsors now have the

opportunity to advertise and add content to our evolving electronic media through the Newslet-

ter, website and Wiki. The Extrusion Division strives to improve our service to you, our mem-

bers, so that your association with the SPE provides you with sustainable business and personal

value.

However, the Extrusion Division needs your help. The Society and our Division depend on

membership, sponsorship and participation. The following are a few of many ways in which

you can participate and "give back" to the Society, the Division and your industry:

1) Renew your membership.

2) Get a new member (see the SPE website for incentives).

3) Attend the Extrusion Division Topcon.

4) Present a paper at the ANTEC 2011.

5) Attend ANTEC.

6) Submit an article or extrusion hint for the Newsletter or WIKI.

7) Become an Extrusion Division sponsor.

8) Participate in Extrusion Division Board of Director activities.

9) Run for a BOD position in the next election cycle.

If you are interested in becoming more active in the Extrusion Division or if you have any com-

ments or suggestions, send me an email at [email protected].

Mark Wetzel

Extrusion Division Chair, 2010-2011.

mailto:[email protected]

Extrusion Division—Society of Plastics Engineers

The Extrusion Division launch of the first Extrusion Wiki in early 2009 has met with broad acceptance.

Board of Directors Member Michelle Curenton was instrumental in developing the Wiki. The Extrusion

Wiki will allow you to search a vast database of information concerning extrusion as well as being able to

submit additional content.

Check it out by clicking here!

Join Us in Boston for ANTEC 2011 May 1-5, 2011

Hynes Convention Center and Boston Marriott Copley Center Hotel

Boston, Massachusetts, USA

Would you like to present a paper at ANTEC? Click here to submit an abstract and a paper!

Abstract and paper deadline: November 19, 2010

http://www.extrusionwiki.com

http://www.extrusionwiki.com/

http://mc.manuscriptcentral.com/speantec


ANALYSIS OF EXTRUSION PROBLEMS WITH

FLOW SIMULATION

John Perdikoulias & Jiri Vlcek

Compuplast International Inc.

This document is meant to demonstrate how process simulation can be used to help determine the causes of

several problems in extrusion. Once the source of the problem is understood, then simulation can also be

used to determine an optimum solution and avoid the wasteful and costly “trial and error” approach.

Problem1: Non-Melted Particles.

The first example is demonstrated with this image that shows some non-melted/non-mixed particles in a sam-

ple of film. These particles are often mistaken for Gels (degraded polymer), but a simple melting test will

confirm if they are Gels or simply non-melted material.

The following image shows a photo of the screw, which was used to produce the above film sample, after it

was removed from the extruder.

The darkening of the screw root in the first half of the screw flights suggests an overheating problem on the

screw root. This may occur if the material cannot progress easily and generates significant friction. To under-

stand this better, one needs to have an understanding of the melting in a single screw extruder. The following

image shows a typical cross section view, of a channel, in the melting region.



FLOW SIMULATION — CONTINUED

A “solid bed” of compressed plastic pellets is formed in the feed section and as a result of frictional

heat at the barrel surface, the material melts and forms a “melt pool” to the rear of the solid bed. The

Solid Bed Ratio is a term defined by the Solid Bed Width divided by the Total Channel Width. As

such, it indicates the fraction of the Channel, along the screw, that is occupied by non-melted mate-

rial. The following image shows the Solid Bed Ratio along the screw length as calculated by the

simulation.

The Solid Bed Ratio is an indicator of how the solid bed melts and is deformed. It can be seen, that

the Solid Bed Ratio starts at 1.0 (100% Solid Material) and begins to reduce around 25% of the way

along the screw. The slope indicates the rate at which the material is melting, or the Solid Bed width

is reducing. However, a little past 40% along the screw, the Solid Bed Ratio starts to increase. This

is due to the compression of the screw root squeezing the solid bed and forcing the width to expand.

This means that the rate of compression is faster than the rate of melting. When the Solid Bed Ratio

goes back to 1.0, then it basically leaves no room for the melted material causing it to penetrate the

Solid Bed and disrupt it.

The following image shows a “snap shot” of the extruder channel as viewed through a glass window

in the barrel.

The “Solid Bed” is the whitish/gray region and the melt is transparent. It can be seen that the Solid

Bed appears to be forming a tear or small gap. This is the initiation of the “Solid Bed Breakup”.

Once this has occurred, the particles become surrounded by the molten polymer and do not experi-

ence the shear against the barrel surface that is required to melt them. As thermal conduction is very

low, these particles then exit the die without having had any mixing or blending with the surrounding

material.




The following image shows the simulation results of a better screw design, with a more gradual compression,

that has a much better Solid Bed Ratio performance.

It can be seen that the Solid Bed Ratio reaches zero well before the end of the screw allowing the material

time to mix and blend properly.

Example 2: Polymer Stagnation and Degradation

The next problem that will be investigated is demonstrated by the image below.

The above image is a cross-section of the material in the transition region from the screen changer to the

adaptor. The darkened material on the curved surface is degraded PVC.

The above flow field was simulated (axisymmetrically) and the Shear Stress is shown in the image below.




The color scale indicates that Blue represents low Shear Stress and Red represents high Shear Stress. Ade-

quate Shear Stress at the polymer/metal interface helps keep the surface clean and avoids polymer stagnation

and degradation. The simulation results show that the location, where the degradation was observed, has very

low wall shear stress.

Once it is understood that it is desirable to maintain an adequate shear stress on the metal surface, a more op-

timum transition channel can be designed, as shown below.

The above simulation results show that a more convex transition results in a higher shear stress along that

wall. This improves the “self-cleaning” characteristics of the flow channel. In fact, a sufficiently high shear

stress should be maintained throughout the entire flow system.




Example 3: Instability in Coextrusion

The final example deals with an Interfacial Instability in a coextruded film sample as shown below.

The film sample is comprised of 3 layers brought together in a feed-block and coextruded through a flat die.

It can be seen that there was intermittent flow in this sample.

The following images show the simulation results on a cross section of the feed-block, where the 3 layers are

brought together.




The black line in the graph (right image) shows the Velocity (left scale) down the center-line of the “white”

layer in the image to the left. The blue line shows the corresponding Elongation Rate (change in velocity)

with its scale being on the right. It can be seen, in the left image, that the middle layer appears to be

“squeezed” after it merges with the other two layers. The corresponding increase in velocity is essentially the

same as stretching or elongating the material.

If the Elongation rate increases too quickly, it can over-stress and essentially “break” the layer. This causes it

to appear intermittently in the product. Once this is phenomena is understood, the simulation can be used to

determine a design that minimizes the amount of Elongation that the material experiences in the merge re-

gion.

The above examples demonstrate how simulation can provide a “window” into the process so that problems

can be better understood. Once the source of the problem is understood, then simulation can be used to help

find a direct solution, quickly and efficiently, and avoid the problem altogether on any future projects.

Extrusion Hints

Remember that for maleic anhydride containing tie resins, you should get them up to 210 deg C minimum in

the extruder even if you want to run them cooler in the die when coextruding with EVOH. If you get them up

to 210 deg C by the metering zone of the screw, this will help to ensure that most of the hydrous form of

maleic is reverted to the anhydride form by the time the material gets into the adapter. In pellet form, MAh

containing polymers absorb moisture from the ambient conditions, and some of the MAh is changed to

maleic acid. If you extrude the polymer at very cool temperatures, you will not open the rings, drive off the

moisture, and convert it back to maleic anhydride. If you are having coextrusion adhesion problems to

EVOH (or Nylon), checking that your extrusion temperature profile through the extruder is proper, is one of

the first things to do.

Scott Marks

DuPont


Bi-Lateral Stress of Bulk Resin Feed

Stephen J. Derezinski, Ph.D.

Extruder Tech, Inc.

Solids conveying in extrusion is often a cause for flow instability and performance issues. Factors typi-

cally considered are the coefficients of friction of the polymer on the barrel surface and on the screw sur-

face. The coefficients are functions of the polymer, the surface condition, the temperature, the rubbing

speed, and the pressure. The pressure (stress) is often assumed to be isotropic. However, bi-lateral pres-

sure in the solid feed is known to exist. A lateral stress that is lower than the primary stress means that the

tractive forces of the friction are developed by some fraction of the primary stress, and this has a signifi-

cant effect on the solids conveying performance of the extruder. Therefore, the level of lateral stress has

just an important role in solids conveying as do the values of friction coefficients.

A laboratory test has been developed1 which is used to determine the bulk lateral stress in polymer sam-

ples as a function of primary stress. The device is shown in Figure 1 as a cutaway drawing. It consists of

a main split cell with a cylindrical chamber to hold the polymer specimen. Pistons at each end are free to

compress the polymer with a press. A load cell is used to record the force on the pistons to determine the

primary stress on the sample.

Figure 1. Cutaway Drawing of the Test Cell1

The lateral force on the specimen is measured by a second load cell. This is possible because the main cell

is split, and the lateral force acts to force the cell apart. The second load cell is shown to resist the separa-

tion and record the lateral force. A bulk value of lateral stress is then obtained from the force recorded by

the second load cell.

The stress values measured are assumed to be bulk average values because they are calculated based on

force divided by applied area. This is desirable because many computational means are based on the bulk

properties of the resin solid feed. Therefore, the lateral stress so measured here is appropriate for applica-


Bi-Lateral Stress of Bulk Resin Feed — Continued

A displacement transducer is used to measure the displacement of the pistons as the polymer is compressed.

The displacement is used to calculate the effective area of the lateral force and the volume of the sample as

compression occurs. The effective area and lateral force are used to calculate the lateral stress as a function

of primary stress. The volume and known mass of the sample also are used to calculate the density of the

sample as a function of primary stress2. Each of these functions is important to analysis of solids conveying,

and each must be included into any comprehensive model of solids conveying.

The values of lateral stress are reduced to a ratio of lateral stress to primary stress. This lateral stress ratio is

then plotted as a function of primary stress to provide a lateral stress ratio function. The function has been

determined to depend on polymer type and polymer form (pellets, powder, etc.) Examples are now presented

for two common polymer types.

Figure 2 shows the lateral stress ratio function of primary stress to be significantly different for three different

polyethylene resins, HDPE, LDPE, and LLDPE. Therefore, lateral stress ratio function is a property of the

polymer type. Also note in Figure 2 that the function is lowest for LLDPE. LLDPE is the most difficult to

extruder of the three polymers shown. LDPE with the greatest lateral stress is not difficult to extrude. This

correlation of extruder performance with lateral stress is typical.

Figure 2. Lateral Stress Ratio Functions for Polyethlyenes1


Bi-Lateral Stress of Bulk Resin Feed — Continued

Figure 3 shows that the lateral stress ratio is unique for the form of the polymer. PET is the resin tested in

powder form and in pellet form. As can be seen, at lower primary stresses the lateral stress ratios are signifi-

cantly different in magnitude and functional shape. As the primary stress is increased the functions logically

approach each other as the resins are compacted closer to becoming a solid mass.

Figure 3. Lateral Stress Ratio Functions for PET Pellets and Powder1

Future work planned for this technology is to study the effect of various amounts of re-cycle on the lateral

stress ratio function. First of all, does the addition re-cycle chips have any effect on the lateral stress ratio? If

so, how much of an effect is created by various amounts of re-cycle chips with the virgin resin bulk feed? Is

there a level of added scrap that will radically affect the extruder performance? The results are planned to be

presented at SPE ANTEC 2011 in Boston, MA in May of 2011.

References

S. J. Derezinski, Measurements of Bilateral Stresses During Compression of Bulk Resin Feed, Conference Proceed-

ings, SPE ANTEC 2010, pp 617-622, 2010.

S. J. Derezinski, Laboratory Measurements of Resin Bulk Density of PET and LDPE, Conference Proceedings, SPE

ANTEC 2009, pp 136-141, 2009.


New Book Review

White, J.L. and Kim, E.K., "Twin Screw Extrusion, Technology

and Principles," 2nd Edition, Hanser, Munich, 2010.

Dr. E.K. Kim and the late Prof. James L. White have just published the

second edition of the "Twin-Screw Extrusion, Technology and Principles"

book through Hanser Publications. The second edition has been expanded

to include many technology advancements since the first edition in 1990.

Technology advancements include those in intermeshing co-rotating twin-

screw machines, both intermeshing and non-intermeshing (tangential)

counter-rotating twin-screw machines, and various continuous mixers such

as the FCM and Buss Kneader. Improved non-isothermal flow numerical

simulation methods are presented along with computer aided design in co-

rotating twin-screw extruders. New application sections were added that

describe reactive extrusion, devolatilization, and dewatering. The book

also provides a rich history of the technologies and the fundamental princi-

ples of operation of twin-screw extruders.

Mark A. Spalding

The Dow Chemical Company

http://www.hanserpublications.com/Products/282-twin-screw-extrusion-2e.aspx


High speed, energy input (HSEI) twin screw extruders are starve fed with the output rate determined by the

feeder(s). Feeders meter solids (pellets/fillers) and/or liquids into the HSEI twin screw extruder for com-

pounding, devolatilization and REX. The extruder screw RPM is independent from the feed rate and is used to

optimize mass transfer efficiencies. Because the pressure gradient is controlled, and zero for much of the

process, materials can be introduced into downstream barrels sections, often by a twin screw side stuffer that

“pushes” materials into the extruder screws. Downstream side stuffing can be beneficial to obtain high filler

loadings, to decrease the abrasive wear in the extruder process section, and when processing shear sensitive

materials. The controlled pressure profile also facilitates venting.

Fig. Example pressure profile in HSEI twin screw extruder

Charlie Martin

Leistriz

Extrusion Hints


Underwater Pelletizing

One of the challenges with underwater pelletizing is having the correct number of die holes for the process to

ensure that all the holes stay open. Obviously you can run with freeze off but this will result in non-uniform

heat distribution in the die plate which in turn can create opportunities for additional problems in pellet uni-

formity. A simple calculation to determine the number of open holes = Throughput (lbs/hour) X 7.6 / (the

weight per pellet(g) X number of knives on the hub X speed of the pelletizer (rpm)). It is a good idea to weigh

at least 100 pellets and take average for the weight per pellet. If more than 10% of the die plate holes are fro-

zen, try increasing throughput or plugging holes.

Tom McHouell

Polymers Center of Excellence

Extrusion Hints

Nylon Processing

Lubricants can, of course, positively influence the extrusion process; however, keep in mind chemical compati-

bility when selecting the proper lube. For example, primary amides, such as erucamide, can actually react in

the extruder with Nylon adversely altering both the lubricant and the polymer chain whereas secondary amides,

such as ethylene bis-stearamide, are more stable allowing the lube to do the job for which it was intended.

Eric Noon

BASF


Mark A. Spalding, Joseph R. Powers, Philip A. Wagner, and Kun Sup Hyun

Dow Plastics, Midland, MI

Abstract

Flow surging in single-screw, plasticating extruders is the variation of the machine’s rate with time, and it generally

leads to higher production costs, lost production, and often higher scrap rates. Flow surging can originate from many

different sources including machine controls, resin feedstock variation, screw geometry, and machine temperature. This

paper will focus on flow surging that originates from improper solids conveying, and it will present experimental data

and corrective action to eliminate or minimize surging.

Introduction Flow surging is defined as the oscillatory change in the output rate of the extruder while maintaining constant set point

conditions. Flow surging can originate from many different sources including improper solids conveying, melting insta-

bilities, and improper control algorithms (1-5). Surging in most case results in lower production rates, higher scrap rates,

material degradation, and higher labor costs.

Previous work documented a severe and random flow surging problem due to improper solids conveying and a solids

obstruction upstream of a spiral dam (4). These surging problems resulted in severe pressure fluctuations at the discharge

of the extruder and thus large rate surges at the die. Rate surges at the die can be estimated from the pressure surges us-

ing the following for flow through a cylindrical restriction (or die) (4):

where n is the power law index, Q1 and P1 are the rate and discharge pressure at condition 1, and Q2 and P2 are the rate

and pressure at condition 2. For example, a 5% variation in the discharge pressure (ÄP = 0.05) for a polymer with a

power law index of 0.3 will cause a 16% change in the instantaneous rate (ÄQ=0.16). An instantaneous rate change of

this magnitude is unacceptable for most processors.

Solids conveying depends on a balance of forwarding forces at the barrel wall and pushing flight and retarding forces at

the screw surface (6). These forces depend mainly on the geometry of the channel and are directly proportional to the

coefficient of dynamic friction for temperatures less than the melting or devitrification temperature and on viscous forces

for higher temperatures (7). Since the coefficient depends on temperature, pressure, and velocity (8), surface temperature

changes for the barrel and screw in the feeding section will strongly affect the performance of the extruder. If the surface

temperatures become too different from the optimal values, flow surging and loss of rate will occur.

Improper design and operation of the melting section of the screw can both lead to extrusion instabilities. For example,

solid bed break up (3) can cause solids to migrate downstream. These solids can wedge into other sections of the screw

and cause the extruder to flow surge (2,4) or cause the extrudate to have periodic changes in temperature. Periodic

changes in discharge temperature will cause some level of flow surging at the die (9).

The goal of this work is to describe the thermal effects of solids conveying on flow surging using two examples for a

high impact polystyrene (HIPS) resin.

Flow Surging in Single-Screw, Plasticating Extruders (Antec 2000 Best Paper)

(1)

(2)

nPQQQQ /1

121 )1(1/)(

121 /)( PPPP


Materials

Rheology, bulk density, coefficient of dynamic friction, and the shear stress at the polymer metal interface were meas-

ured for the HIPS resin. The rheology and bulk density of the resin were essentially the same as that previously reported

for a slightly different HIPS resin (10,11). The shear stress at the polymer-metal interface is reported here rather than the

dynamic friction since friction is only defined for solid-state processes, while the stress can be described from ambient

temperatures up to processing temperatures. The shear stress at the interface (12) for HIPS resin is shown by Figure 1 at

a pressure of 0.69 MPa. As indicated by this figure, the shear stress was nearly constant from ambient temperature up to

about 110°C, increased to a maximum stress near 150°C, and then decreased as the temperature was increased further.

Optima! performance of the solids conveying section for this resin would be such that the forwarding forces are maxi-

mized with metal surface temperatures near 150°C where the stress is a maximum, and the retarding forces minimized

with metal surface temperatures of 110°C or lower. Thus, optimal solids conveying for HIPS would occur with a feed

zone barrel inner surface temperature near 150°C and a screw surface temperature in the feed section no higher than

110°C. In practice, screw temperatures less than 90 or 100°C are preferred such that melting of the resin does not occur

if an emergency shutdown should occur. For the solid state temperature region, the shear stress at the interface can be

converted to the coefficient of dynamic friction by the following (12):

(3)

where f is the coefficient of dynamic friction, is the shear stress at the polymer metal interface, and P is the pressure

(0.69 MPa).

Screw Temperature Control

A severe and random flow surging problem limited the production rate for a large-diameter, two-stage, vented extruder.

if it were not for a gear pump positioned between the extruder and die, this extrusion line would not be operable. The

surging did, however, limit the output of the line to about 60% of its potential rate. The maximum potential rate is the

rate that the extruder can run at high screw speed and with proper operation. In order to diagnose the problem, a data ac-

quisition system was temporarily connected to the extrusion panel. The screw was single-flighted and typical of what is

used for HIPS resins. Screw temperature control was accomplished by flowing cooling water through a rotary union into

and out of a hole cut into the feed end of the screw (13). This hole extended to about 4 diameters into the feed section.

Pressure sensors were positioned in the barrel wall at the end of the first-stage transition section (P1), at the end of the

first-stage metering section just before the vent (P2), and at the discharge. An additional pressure sensor was positioned

at the inlet (suction side) to the gear pump. A temperature sensor was positioned in the transfer line upstream of the gear

pump to measure extrudate temperature. A commercial control scheme adjusted the screw speed to maintain a constant

pressure of 9 MPa to the inlet of the gear pump. The gear pump was operated at constant speed in order to maintain a

constant flow rate of material to the downstream equipment.

Steady-state operation of the extruder is shown by the first 400 minutes in Figures 2, 3, and 4. The data for these figures

were from the same experimental run. At these conditions, the extruder was operating at about 70% of its potential maxi-

mum rate and the screw speed varied only about ± 2 rpm. The barrel pressure at the end of the first-stage transition sec-

tion, P1, had variations of about ± 3 MPa about the average pressure. This pressure variation was considerably higher

than was expected and suggests that the extruder, although running stable, was on the verge of unstable operation. Some

of the variation was due to the movement of the flight tip past the sensor. Barrel zone temperatures tracked the set point

values and were stable. Calculations indicated that the first stage was full of resin.

Pf /

Flow Surging in Single-Screw, Plasticating Extruders

— Continued


At about 410 minutes into the run, the extruder started to operate unstably, as indicated by Figures 2, 3, and 4. The proc-

essing change that caused the extruder to go from a stable operation to an unstable one is not known, but it could have

been due to minor changes in the bulk density of the feedstock or cooling water fluctuations. As indicated by these fig-

ures, the event started when the P1 pressure decreased slightly, causing the rate and the P2 pressure to decrease. This

decreased pressure transmitted down the extrusion system, eventually decreasing the pressure at the inlet to the gear

pump. To correct for the lower pressure, the controller on the gear pump increased the speed of me screw from 100 to

about 160 rpm. Next the Pt pressure increased due to the higher screw speed and higher flow rate, as indicated by Figure

3. As the pressure increased at the gear pump inlet, the gear pump controller decreased the screw speed back to 100 rpm,

causing the extruder to flow surge. Flow surging caused the screw speed controller to oscillate about once every 25 min-

utes. As indicated by Figure 2, the screw speed controller was able to provide a relatively stable pressure to the pump

inlet, allowing the process to run at reduced rates. The barrel zone temperatures, as indicated by Figure 4, were ex-

tremely oscillatory.

As indicated by Figure 2, the P1 pressure was considerably lower during the period of unstable operation. This result

indicates that the cause of the problem originated in the first stage of the screw before the metering section. At a screw

speed of 160 rpm, the extruder was capable of a rate of at least 50% greater than the actual rate. This lower than expected

rate indicated that the first-stage metering section was operating only partially filled. The most likely reason for a starved

metering section was poor solids conveying from the feed section to the transition section. This poor solids conveying

was likely due to improper temperature control of the metal surfaces in the feed section of the extruder and screw. Barrel

feed zone heaters and controllers were examined and determined to be operating properly at set point temperatures typi-

cally used for HIPS. Based on this information, the investigation was focused on screw temperature control.

The effect of internal screw cooling was determined during a period when the extruder was operating stably. For this

stable extrusion period, cooling water was flowing to the screw-cooling lance and the extruder was operating stably and

properly at a rate of 70% of its potential maximum. The metal surface temperatures of the pipes used to flow water into

and out of the screw were measured at 29 and 37°C, respectively. At about 28 minutes into the run, the cooling water

flow to the screw was turned off, as indicated by Figures 5 and 6. At about 30 minutes, the pressure at the end of the first

-stage transition section, P1, started to decrease as shown by Figure 6, indicating that solids conveying was significantly

reduced. Like before, the reduced solids flow caused the downstream pressures to decrease, and ultimately to cause the

extruder to flow surge. At about 36 minutes into the run, cooling water flow was turned on and within about 4 minutes

the extruder operation became stable, as indicated by the Figures 5 and 6. The surface temperature of the pipe for water

flow out of the screw was measured at 81°C just after the cooling water was turned on; a temperature change of 44°C.

Previous research on HWS (14) has shown that solids conveying becomes difficult or unstable at screw temperature of

about 160°C and higher. The temperature of the screw surface was unknown, but it likely increased by at least 44°C and

possibly approached 160°C.

Based on the above data, the cause of the extrusion instability was identified as high screw surface temperatures in the

feed section. These high surface temperatures caused the coefficients of dynamic friction to increase, increasing the re-

tarding forces on the solids at the screw surface. Since solids conveying depends on a combination of forwarding forces

at the barrel wall and pushing flight and retarding forces at the screw root and trailing flight, an increase in the retarding

forces will cause a reduction in the solids conveying rate. The instability appeared to be random due to the complicated

interactions of cooling water flow rate and temperature, and changes in bulk density of the feedstock.

To increase the operating window for stable extrusion, the cooling level to the feed section of the screw was increased by

lengthening the cooling hole in the screw. The cooling hole length was increased from 4 diameters into the flighted sec-

tion to 7 diameters; i.e., up to the end of the feed section. After the screw modification, the extruder has not experienced

instabilities of this type and the rate was increased to 100% of its maximum potential rate.


— Continued


Feed Casing Temperature Control

On a different occasion, the same extruder started to flow surge, but with a slightly different frequency, as shown by Fig-

ures 7 and 8. As indicated by these figures, there were short time periods when the discharge pressure and screw speed

were stable and the motor current was high. During these periods, the extruder was operating well but at a reduced pro-

duction rate. During periods of unstable operation, the motor current decreased by about 20%. the screw speed increased,

and the discharge pressure became extremely oscillatory. Like the previous case, as the motor current decreased solids

conveying decreased, causing the controller to increase the speed of the screw.

During the trial, the feed casing to the extruder had an outside surface temperature of about 80°C. Although not meas-

ured, the inside cylinder wall of the feed casing for first 1.5 diameters downstream of the feed opening was considerably

hotter. These higher temperatures were caused by a combination of frictional heating of the solids on the wall and also

by conduction from the first heated zone of the barrel. It is estimated that temperatures as high as 170°C occurred in the

feed casing. Optimal solids conveying will occur when the stress at the polymer-metal interface at the barrel is a maxi-

mum. This maximum stress occurs for HIPS at a surface temperature near 150°C, as indicated by Figure 1. Surface tem-

peratures higher than 150°C in the feed section will reduce conveying and lead to starving of the screw channels and ulti-

mately flow surging. Corrosion inside of the cooling channels of the feed casing prevented the flow of cooling water.

Cleaning the cooling channels and adding a larger cooling water recirculation pump reduced the temperature of the feed

casing and eliminated the flow surging problem.

Discussion

The examples presented show the effect of improper surface temperatures in the feed section of a plasticating extruder.

Acceptable solids conveying will occur when the melting and metering sections are full of resin and under pressure; i.e.,

the first-stage metering section must control the rate. If the solids conveying section is rate controlling as in the two ex-

amples, then downstream sections of the screw channel will be starved leading to low rates, material degradation, and

flow surging.

Optimal conveying forces occur when the forwarding forces are maximized and the retarding forces are minimized. The

forwarding forces occur at the barrel-polymer interface and at the pushing flight, while the retarding forces occur at the

screw root and the trailing side of the screw (6). These forces originate from solid-state friction at low temperature (7)

and by viscous forces at higher temperatures (8,11). Both types of forces are proportional to shear stress between the

rubbing polymer and the metal surface. For HIPS, this shear stress is shown by Figure 1. As indicated by this figure, the

shear stress is relatively constant at temperatures less than 100°C, has a maximum stress at about 150°C, and then de-

creases with increasing temperature. In order to maximize the forwarding force at the barrel surface, it is obvious from

Figure 1 that the barrel surface temperature in the feed section should be about 150°C. In practice, an axial temperature

gradient will exist between the water cooled feed casing and the first heated zone of the extruder barrel, and temperature

sensors are generally not capable of measuring the surface temperature of the barrel accurately (15). Thus, an experimen-

tally determined first zone temperature in the range of 150 to l70°C and a feed casing temperature around 35 to 45°C is

acceptable. For the screw, both forwarding and retarding forces occur, yet the temperature of the screw surfaces must be

controlled to the same temperature. Experience has shown that optimal screw temperatures for HIPS are those less than

90°C. For a temperature range of ambient up to 90°C, the retarding forces are essential the same, and by reducing the

temperature below the glass transition temperature, devitrification (melting) will not occur on the screw root during an

emergency shutdown.


— Continued

Conclusions

Transient process data clearly show the effects of feed zone temperature control on the extrusion performance of HIPS.

Optimal feed zone barrel and screw temperature control must ensure that the solids conveying section delivers adequate

resin to keep the downstream sections of the first stage full. The optimal temperatures maximized the solids conveying

forwarding forces and minimized the retarding forces.

References

1. Z. Tadmor and I. Klein, "Engineering Principles of Plasticating Extrusion," Van Nostrand Reinhold Co., New York,

1970.

2. I. Klein, SPE J., 28, 47 (1972).

3. R.T. Fenner, A.P.D. Cox, D.P. Isherwood, SPE ANTEC Tech. Papers, 24,494 (1978).

4. K.S. Hyun and M.A. Spalding, Adv. Polym. Tech., 15, 29 (1996).

5. K.S. Hyun, M.A. Spalding, and J. Powers, Plast. Eng., 52, 4, 33 (1996).

6. K.S. Hyun and M.A. Spalding, SPE ANTEC Tech. Papers, 43, 211 (1997).

7. CI. Chung, SPE J., 26, 32 (1970).

8. M.A. Spalding, D.E. Kirkpatrick, and K.S. Hyun, Polym. Eng. Sci., 33,427 (1993).

9. J.R. Thompson, SPE ANTEC Tech. Papers, 40, 288 (1994).

10. J. Dooley, K.S. Hyun, and K. Hughes, Polym. Eng. Sci., 38,1060 (1998).

11. K.S. Hyun and M.A. Spalding, Polym. Eng. Sci., 30, 571 (1990).

12. M.A. Spalding, K.S. Hyun, and B.R. Cohen, SPE ANTEC Tech. Papers, 43, 202 (1997).

13.S.R. Jenkins, J.R. Powers, K.S. Hyun, and J.A. Naumovitz, J. Plastic Film and Sheeting, 6, 90 (1990).

14. M. Mizoguchi, Japan Steel Works Tech. News, 11, 1 (1975).

15. T.W. McCullough and M.A. Spalding, J. Reinforced Plast. Comp., 16, 1622 (1997).


— Continued



— Continued

0

0 .1

0 .2

0 .3

0 .4

0 .5

0 50 100 150 200 250 300

T e mpe ratu re , C

Sh

ea

r S

tre

ss

, M

Pa 7 .6 c m /s

1 5 .2 c m /s

3 0 .5 c m /s

Figure 1. Shear stress between HIPS and a metal surface at 0.69 MPa and as a function of sliding velocity.

0

5

10

15

20

25

0 200 400 600 800 1000

T ime , minu te s

Pre

ss

ure

, M

Pa

0

200

400

600

800

1000

1200

1400

Mo

tor c

urre

nt,

A

Current

P1

P2

Pump Inlet

Extruder Discharge

Current

P1

Figure 2. Barrel, discharge, and pump inlet pressures and motor current for stable and unstable extrusion for a large-

diameter extruder running HIPS.

80

120

160

200

240

280

0 200 400 600 800 1000

T ime , min u te s

Sc

re

w s

pe

ed

, rp

m

0

200

400

600

800

1000

1200

1400

Mo

tor c

urre

nt,

A

Screw Speed

Current

Figure 3. Screw speed and motor current for a large-diameter extruder running stable and unstable.



— Continued

50

100

150

200

250

300

0 200 400 600 800 1000

T ime , minu te s

Te

mp

era

ture

, C

80

120

160

200

240

280

320

Sc

re

w s

pe

ed

, rp

m

Screw Speed

Extrudate

T8T9

T7

T2,T3

T1,T4,

T5,T6

Figure 4. Screw speed, extrudate temperature, and barrel zone temperatures for a large-diameter extruder running stable and

unstable.

0

2 0 0

4 0 0

6 0 0

8 0 0

1 0 0 0

1 2 0 0

1 4 0 0

0 2 0 4 0 6 0

T im e , m in u tes

Mo

tor c

urre

nt,

A

8 0

1 2 0

1 6 0

2 0 0

2 4 0

2 8 0

3 2 0

3 6 0

Sc

re

w s

pe

ed

, rp

m

Screw Speed

Current

Screw

cooling

off

Screw

cooling

on

Figure 5. Screw speed and motor current for the screw cooling experiment.

0

10

20

30

40

0 20 40 60

Time, minutes

Pre

ssure

, MP

a

0

40

80

120

160

200

Scr

ew

speed, r

pm

Screw speed

P1

Extruder discharge

Screw

cooling

off

Screw

cooling

on

Figure 6. Screw speed, pressure at the entry to the first-stage meter (P1), and discharge pressure for the screw cooling experi-

ment.



— Continued

40

80

120

160

200

240

280

0 20 40 60

T im e , m in u tes

Sc

rew

sp

ee

d,

rpm

0

200

400

600

800

1000

1200

1400

Mo

tor

cu

rre

nt,

A

Current

Screw Speed

Figure 7. Screw speed and motor current for a large-diameter extruder with a feed casing that was too hot.

0

5

10

15

20

25

0 20 40 60

T im e , m in u tes

Pre

ss

ure

, M

Pa

0

200

400

600

800

1000

1200

1400

Mo

tor

cu

rre

nt,

A

Current

Pressure

Figure 8. Discharge pressure and motor current for a large-diameter extruder with a feed casing that was too hot.

Interfacial instabilities in coex films can often give the appearance of gels. These “gels” are actually one polymer

flowing across the interface into the other polymer layer due to localized rheology mismatches. The phenomena is

most common at chemically bonded interfaces between tie-resins and EVOH or Nylon but can also occur at non

bonded interfaces with significantly different viscosities mismatches. This type of interfacial instability is difficult

to correct by changing processing conditions alone and usually require material changes. For best results, estimate

shear rates in the die and choose materials with similar viscosities at those shear rates. When small viscosity mis-

matches are unavoidable, it is best to run the lower viscosity material closer to the channel wall.

Dan Ward

Nova Chemicals

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2010 SPE Extrusion Division Board

Andersen, Paul G., Ph.D. Coperion Corporation [email protected]

Anzini, David Zip-Pak [email protected]

Biesenberger, Jeff Advanced Drainage Systems [email protected]

Bigio, David I., Ph.D. University of Maryland [email protected]

Campbell, Gregory A., Ph.D. Clarkson University [email protected]

Christiano, John Davis-Standard LLC [email protected]

Curenton, Michelle Solo Cup [email protected]

Cykana, Daniel Bemis Manufacturing [email protected]

Derezinski, Stephen J., Ph.D. Extruder Tech, Inc. [email protected]

Golba, Joseph C., Jr., Ph.D. PolyOne [email protected]

Gould, Russell J. RG Associates [email protected]

Gupta, Mahesh Ph.D Michigan Technological University [email protected]

Karszes, William Plastics Associates [email protected]

Larson, Keith ACS Colortronics [email protected]

Martin, Charlie Leistriz [email protected]

Mchouell, Tom Polymer Center for Excellence [email protected]

Morris, Barry A., Ph.D. DuPont [email protected]

Mount, Eldridge, III, Ph.D. EMMOUNT Technologies [email protected]

Myers, Jeffery Barr Incorporated [email protected]

Oliver, Gary Extrusion Dies Industries [email protected]

Perdikoulias, John, Ph.D. Compuplast Canada [email protected]

Puhalla, Mike Milacron Incorporated [email protected]

Schick, Steven F. Teel Plastics Inc. [email protected]

Schildknecht, Helmut List Incorporated [email protected]

Spalding, Mark, Ph.D. The Dow Chemical Company [email protected]

Wagner, John R., Jr. Crescent Associates [email protected]

Wetzel, Mark DuPont [email protected]

Womer, Tim Xaloy [email protected]

Xiao, Karen, Ph.D. Brampton Engineering Inc. [email protected]







mailto:steve@extrudertd



















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Society of Plastics

Engineers

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promote the scientific and engineering education and

knowledge relating to the extrusion of plastics.”

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Industry Experts at the Extrusion Division Website:

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contact:

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volume 35 winter 2010 message from the extrusion division ... · extrusion problems with flow...

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