US Army Corps

of Engineers

Construction Engineering

Research Laboratories

Development of Recycled Polymer Blends

for Thermal Spray Applications

by

J. Brogan, C.C. Berndt, S. Sampath, H. Herman, and S.A. Drozdz

Approved for public release; distribution is unlimited.

A Corps/Industry Partnership To Advance

Construction Productivity and Reduce Costs

USACERL Technical Report 98/14

December 1997

CONSTRUCTION PRODUCTIVITY ADVANCEMENT

RESEARCH (CPAR) PROGRAM
USACERL TR 98/14 1

SF 298

A proposed rule to implement Section 183(e) of the Clean Air Act amendments

(1990) will further restrict the volatile organic compound (VOC) content of

rust-preventing coatings. Thermal spray polymers represent one option for

protecting steel structures while complying with the new VOC restrictions.

Thermal spray technology can apply polymer coatings to many different substrate

materials at various thicknesses and under wide range of ambient conditions.

Potential benefits of thermal spray polymers could be extended by blending

post-consumer commingled polymers (PCCP) into coating feedstock materials;

thermal spray polymers have the potential both to lower material costs and reduce

the volume of the solid waste stream. This research investigated and demonstrated

high-molecular weight and commingled/postconsumer recycled polymer blends as

low-VOC thermal spray coatings.

It is concluded that the incorporation of PCCP into virgin thermal spray polymer

is associated with some loss of coating performance. Also, cost savings resulting

from blending 25 percent PCCP with 75 percent virgin polymer are insignificant

compared to the total cost of a recoating job using other VOCcompliant systems.

Therefore, the blends investigated here probably will not have significant

commercial usefulness, and they are not recommended for corrosion protection of

steel structures.
2 USACERL TR 98/14

Foreword

This study was conducted for Headquarters, U.S. Army Corps of Engineers under

Construction Productivity Advancement Research (CPAR) Work Unit 6L2,

“Development of Innovative Hi-Molecular Weight and Commingled/Recycled

Polymer Blends for Thermal Spray Applications.” The technical monitors were D.

Casapula, CECW-EE, and A. Chandry, CECW-OM.

This CPAR work unit was coordinated through the Materials Science and Technology Division (FL-M) of the Facilities Technology Laboratory (FL), U.S. Army

Construction Engineering Research Laboratories (USACERL). The work was

conducted through a CPAR Cooperative Research and Development Agreement

(CRDA) between USACERL and The State University of New York (SUNY) at

Stony Brook. The USACERL Principal Investigators were Richard G. Lampo and

Susan A. Drozdz, CECER-FL-M. The SUNY Principal Investigator was Dr.

Christopher C. Berndt. Dr. Ilker R. Adiguzel is Acting Chief, CECER-FL-M, and

Donald F. Fournier is Acting Operations Chief, CECER-FL. The USACERL

technical editor was Gordon L. Cohen, Technical Information Team.

COL James A. Walter is the Commander of USACERL and Dr. Michael J. O’Connor

is the Director.
USACERL TR 98/14 3

Contents

SF 298 ................................................................ 1

Foreword .............................................................. 2

List of Figures and Tables ................................................ 4

1 Introduction ....................................................... 7

Background ........................................................ 7

Objective .......................................................... 8

Approach ......................................................... 9

2 Technology and Experimental Techniques ............................. 10

Thermal Spray Processing of Polymers .................................. 10

Thermal Spray Torches .............................................. 10

Polymer Powder Analysis ............................................ 11

Testing and Analysis Technologies ..................................... 12

3 Experimental Results .............................................. 19

Flame Spray Processing of Virgin Polymers .............................. 19

Flame Spray Processing of Recycled Polymer Blends ....................... 26

Mechanical Properties of Polymer Materials and Blends ..................... 29

EMAA Adhesion to Steel ............................................. 32

4 Field Application .................................................. 62

5 Cost Considerations ............................................... 64

6 Conclusions, Recommendations, and Commercialization ................. 65

Conclusions ...................................................... 65

Recommendations ................................................. 66

Commercialization and Technology Transfer .............................. 66

References ........................................................... 69

Appendix: Equipment and Instrumentation Vendors ............................ 71

Distribution
4 USACERL TR 98/14

List of Figures and Tables

Figures

1 Schematic of combustion spray torch ........................... 16

2 The effect of gas mixture upon (a) flame temperature and (b)

flame length .............................................. 17

3 Construction of peel test jig ................................... 18

4 Peeled polymer coating from abrasive-blasted substrate ............. 18

5 Ethylene methacrylic acid copolymer powder ..................... 36

6 Cross-section of coating with low splat elongation ratios ............. 38

7 Cross-section of coating with high splat elongation ratios ............ 39

8 Distribution of elongation ratios for the microstructural extremes .................................................. 39

9 Relationship between elongation ratio and surface roughness

........................................................ 40

10 Relationship between surface roughness and deposit temperature ..................................................... 40

11 Relationship between splat elongation ratio and deposit

temperature .............................................. 41

12 Morphology of post-consumer commingled plastic after

grinding ................................................. 41

13 Cross-section of post-consumer commingled polymer coating

........................................................ 42

14 Surface of as-sprayed PCCP coatings from dry compounded

recycled plastic ............................................ 43
USACERL TR 98/14 5

15 Surfaces of weathered polymer coatings ........................ 44

16 Cracking mechanisms for recycled post-consumer commingled

polymer blends: (a) interfacial debonding, (b) crack intersection, (c) inter-particle cracking and (d) radial cracking ............... 45

17 Pseudo crack density of exposed coatings ....................... 45

18 Average crack length for exposed coatings ....................... 46

19 Tensile adhesion strengths for PCCP coating blends ............... 46

20 Stress-strain response of PF111 when pigmented, clear, and

quenched ................................................ 48

21 Selected mechanical properties for PF111: a) modulus, b)

tensile, c) elongation to break, and d) toughness .................. 49

22 Stress-strain response for PF113 when pigmented, clear, and

quenched ................................................ 50

23 Selected mechanical properties for PF113: (a) modulus, (b)

tensile, (c) elongation to break, and (d) toughness ................. 51

24 Stress-strain response for PCCP blends; volume percent

PCCP noted on plot ........................................ 52

25 Mechanical properties of PCCP blends: (a) yield strength, (b)

tensile strength, (c) secant modulus, and (d) elongation to

break ................................................... 53

26 Summary of peel strength obtained for varous steel surfaces ......... 55

27 Load-displacement plot for PF111 on polished steel ................ 55

28 Polymer fracture surface from polished steel depicting bands

at (a) low and (b) higher magnification .......................... 56

29 Load-displacement plot for PF111 on polished steel ................ 57

30 Peeling fracture surfaces from abrasive-blasted steel (a)

PF111 and (b) PF113 ....................................... 58
6 USACERL TR 98/14

31 EMAA peeled from abrasive-blasted steel ........................ 59

32 (a) Load displacement plot for as-received steel surface and

(b) SEM micrograph of polymer coating surface ................... 60

33 Proposed interfacial interactions between EMAA copolymer

and steel oxide ............................................ 61

34 Field demonstration at the Triborough Bridge: (a) preheating

the steel panel and (b) deposition of recycled polymer blend .......... 63

Tables

1 Thermal-spray parameters for design of experiments study .......... 37

2 2

5-1 design for EMAA copolymer ............................... 37

3 Statistically significant parameters* ............................. 38

4 Bulk properties of individual polymer constituents .................. 42

5 Summary of spray-formed EMAA mechanical properties ............ 47

6 Mechanical properties of melt-compounded PCCP with EMAA

copolymer ................................................ 52
USACERL TR 98/14 7

1 Introduction

Background

Protecting the infrastructure from the effects of corrosion is an increasing concern

for the Army and the nation in general [1]. The average age of many bridges, locks,

and dams exceeds 40 years, so maintenance is essential to ensure their reliability

and safety of the populace. Choosing a method of corrosion protection is essentially

a materials selection process, in which one chooses a protective coating system and

perhaps a cathodic protection scheme [2, 3, 4]. Factors that influence this choice

include the forms of corrosion encountered, exposure conditions (which range from

mild atmospheric to industrial or marine), the type of structure and its condition,

geographic location, and budget.

A proposed rule to implement Section 183(e) of the Clean Air Act, as amended in

1990, will limit the volatile organic compound (VOC) content of rust-preventing

coatings to 400 grams per liter (61 FR 3279–32746; 40 CFR 59). When this rule

takes effect on 1 January 1998, VOC content will be a major factor in the selection

of protective coatings, and some coating systems used today (i.e., high-VOC alkyds,

urethanes, and epoxies) will be prohibited [7]. Alternative technologies and

non-solvent based coating materials and application technologies are currently

being researched. Thermal sprayed thermoplastic is one viable option for meeting

the construction industry's needs while complying with environmental restrictions;

it is a 100 percent solids process that releases minimal VOCs [8]. Relatively thin

(0.127 mm [0.005 in.]) to thick (6.35 mm [0.250 in.]) polymer coatings can be

applied onto a wide variety of substrate materials to produce protective barrier

coatings [5, 6]. Thermoplastic spraying has potential as an alternative to paint and

other solvent-borne formulations.

In 1994 the U.S. Army Construction Engineering Research Laboratories

(USACERL) published a technical report on an evaluation of flame-sprayed polymer

coatings for civil works navigation structures [9]. The coatings were not

recommended for use at that time because of their relatively poor performance in

laboratory and field tests. The coatings were tested in immersion in fresh and salt

waters in addition to atmospheric exposure. An ethylene-acrylic acid copolymer

provided the best performance in all exposures, but its expected service life of 5
8 USACERL TR 98/14

years in immersion service compares poorly to the 15 to 20 years for currently used

vinyl systems. Adequate performance and an 8 to 10 year service life was

estimated for atmospheric weathering — much shorter than the 20 to 30 years of

protection provided by standard Army Corps of Engineers systems.

Despite its unremarkable performance in these tests, thermal spraying of polymers

offers a number of compelling potential benefits for the U.S. construction industry.

However, successful applications will require the definition of structural types,

substrates, and environmental conditions for which thermal-sprayed polymers can

be used. Polymer spraying is a one-coat process; the coating acts as both primer

and topcoat. The coating requires no additional cure times (unlike the traditional

multicoat painting processes) [6]. Thermal spraying is adaptable to structures of

any size. Unlike electrostatic powder coatings, this process can coat nonconductive

components (e.g., concrete, plastic, fiberglass) and can be used on heat-sensitive

materials (e.g., wood and wood products). Thermoplastic materials can always be

remelted and will fuse with new material in the molten state, so it is possible that

these coatings could be repaired by simply remelting or applying additional material to the damaged area. In addition, a report on thermal spray powder coatings

[10] shows that because these coatings do not rely on solvent evaporation for film

deposition, polymers such as ethylene methacrylic acid copolymer (EMAA) can be

applied in high humidity as well as in temperatures below freezing — conditions

under which traditional paint systems often may not successfully be applied.

The potential benefits of polymer thermal spraying could be extended further by the

incorporation of post-consumer (recycled) polymers into the coating feedstock.

However, the use of commingled recycled polymer materials into thermal spray

polymer coatings has not been reported in the literature. Commingled polymer,

also referred to as "curbside tailings," is a mixture of many types of thermoplastics,

and often contains paper, aluminum, and glass particulate wastes. This particular

blend of plastics is the remaining yield after high-value polymer components such

as polyethylene terphthalate (PET) and high-density polyethylene (HDPE) are

separated out. It is therefore of considerable interest to determine the viability of

using recycled and post-consumer commingled polymer in the form of coatings to

reduce cost as well as reduce the amount of waste disposal.

Objective

The objective of this program was to develop and demonstrate innovative highmolecular weight and commingled/post-consumer recycled polymer blends as lowVOC coatings for thermal spray application.
USACERL TR 98/14 9

Approach

Thermal sprayed polymeric coatings are generally applied by melting polymer

powder in a combustion flame or plasma stream. The material carried by the

heated gas is deposited as "splats" (or droplets of molten plastic) on the substrate

surface, where it fuses and consolidates into a coating film. Thermal spray

equipment, designed exclusively for polyethylene-based polymers (as marketed by

Plastic Flamecoat Systems, Big Spring, TX), allows simultaneous thermoplastic

deposition and fusion. Ethylene methacrylic acid copolymer (EMAA) is the most

commonly used polymer in the field, but little structure-property data are reported

in the open literature. In this study, commercially available equipment and EMAA

copolymer powder were used for laboratory and field applications.

A design of experiments approach was used to determine how the thermoplastic

microstructure is affected by variations in process parameters such as substrate and

coating temperature, flame setting, and standoff distance. Mechanical properties

were then assessed to determine optimum processing conditions.

The research objective included the characterization of coatings produced from

post-consumer commingled polymer (PCCP). The objective is complicated by the

fact that incompatible polymers in the commingled mix would melt and coalesce

during thermal spraying. To address this problem, PCCP was blended in various

percentages with virgin EMAA copolymer for improved performance. The

properties of commingled polymer coatings were compared to virgin EMAA coatings

to assess whether the use of these waste materials is accompanied by a

degradation in mechanical properties.

Commingled polymer particulates experience a variety of thermal histories

depending upon their individual thermal diffusivity, particle size, molecular weight,

residence time, and position within the flame. These factors in turn influence the

degree of polymer flattening as well as the rate of cooling and solidification.

Microstructural features (splat elongation ratio), polymer structure and chemistry

(oxidation, crystallinity, molecular weight), adhesion, and mechanical properties

were analyzed in relation to selected processing parameters. Specifically, analyses

were conducted to determine such structure-property relationships for EMAA

copolymer, to understand the mechanisms affecting adhesion to steel surfaces, and

to characterize PCCP-based coatings.
10 USACERL TR 98/14

2 Technology and Experimental Techniques

Thermal Spray Processing of Polymers

In the thermal spray polymer coating application process, the polymer powder is

axially fed into the combustion zone via a carrier gas. The polymer particles are

propelled through the flame, where they melt into droplets and are transported to

a preheated substrate. As the molten particles hit the substrate, the molten

particles deform and solidify, creating an interlaced network of “splats.” Porosity

within the coating depends upon the coating temperature, nature of the substrate

surface, and the polymer melt-viscosity. The thickness of the coating is governed

by droplet velocity, rate of spray gun movement, and number of passes across the

substrate.

Polymer powders are specified by their chemistry, morphology, molecular weight

distribution and associated melt-flow index, and particle size distribution. Spray

parameters must be selected to accommodate each particular polymer formulation.

A larger thermal input is required to melt larger particles and higher molecular

weight organic molecules. A large particle size distribution or molecular weight

distribution may facilitate the formation of numerous heterogeneities within the

coating microstructure, creating voids, a range of splat aspect ratios, and degraded

material. A “processing window” exists for each polymer, where poor particle

coalescence defines the lower limit and pyrolysis defines the upper limit. It is also

important to note that, even within the processing window, the polymer may also

cross-link, undergo chain scission, or oxidize.

Thermal Spray Torches

Plastic Flamecoat 200 Torch

The PFS 200 gun (marketed by Plastic Flamecoat Systems [PFS], Big Spring, TX)

is the prototype torch initially used in the course of this research [14]. This

particular torch uses compressed air as both the oxidant for combustion and as the

carrier gas to propel the powder through the flame. Recommended specifications

include 283 l/min (10 cfm) of air at 414 kPa (60 psi) as well as 5–15 pounds of
USACERL TR 98/14 11

*

Figures and tables are presented at the end of the chapter in which they are first referenced.

propane pressure. Powder is conveyed via compressed from a gravity-fed hopper.

Flame air, propane flow rate, and powder flow rate can independently be controlled.

The larger torch PF400 has a 4-in. (10 cm) diameter nozzle in comparison to the

standard two-inch design.

Plastic Flamecoat Powder Pistol 124

The Powder Pistol 124, the company’s second-generation combustion-spray gun, is

connected to a fluidized bed powder feeder (Figure 1).*

 Propane and compressed air

are used as the fuel and oxidant for the combustion process. The air supply

pressure ranges from 621-827 kPa (90-120 psi) and generally is consumed at 198

l/min (7 cfm). The air supply should be free of oil and moisture. The compressed

air supply provides air for the propane/air mixture, provides air for powder

transport through the flame, and provides air to fluidize the powder. The powder

feeder can hold a maximum of 10 lb (4.5 kg) EMAA powder. In addition, the

maximum application rate is approximately 8 lb/hr. The recommended propane

supply pressure ranges from 90-117 kPa (13-17 psi) and generally is consumed at

1.5 lb/hr (0.68 kg/h). The flame air, fluidized bed air, powder feed rate, and propane

flow rate can be independently controlled. It should be noted that a retrofit is

available, increasing polymer application rates to 12.5 lb/h (5.7 kg/h).

Polymer Powder Analysis

EMAA powder was received from Plastic Flamecoat Systems. Pigments are incorporated using dry compounding or melt compounding procedures. Proprietary additives, including UV-light stabilizers, antioxidants, and flow enhancers are also

blended to optimize coating properties and ease of processing. The base-polymer is

cryogenically ground to -70 mesh (< 212 micrometers), resulting in a particle

morphology is angular and faceted. The particle becomes more spherical in shape

during the melt-compounding operation because the particle is heated above its

crystalline melting point. Consequently, it is more difficult to fluidize or feed the

unpigmented powder due to its angular morphology.

Two EMAA powders were used in this research: PF111 and PF113. Powder size

distributions were obtained using Micro Trac Powder Analysis. The average size,

assuming a spherical shape factor was 140 micrometers with a standard deviation

of 70 micrometers. Ninety percent of the powder was also less than 220

micrometers. These powders were chosen on the basis of availability and the
12 USACERL TR 98/14

discernible difference in melt-flow characteristics. As previously noted, EMAA

copolymer has the best balance of performance properties (i.e., adhesion and

permeability).

Testing and Analysis Technologies

Flame Temperature

Flame temperatures were measured with a thermocouple at 5 cm (approximately

2 in.) intervals from the nozzle. The highest flame temperature was approximately

1300 EC (2370 EF). Temperatures and flame lengths were dependent upon the

flame stoichiometry. Figure 2 summarizes the effects of gas mixture upon

temperature and flame length. The higher the air-to-fuel ratio, the higher the

flame temperature at distances close to the nozzle (5-10 cm [2-4 in.]). Peak

temperatures for fuel-rich flames were observed further downstream as reflected

in the data at 45 cm (approximately 18 in.). Fuel-rich flames were also much longer

in length, approximately 50 cm (20 in.) compared to approximately 12 cm (5 in.) for

a high air-to-fuel ratio.

Melt Flow Indexer

The melt flow indexer is an industrial device used to assess the molecular weight

of polyethylene. A melt flow indexer, also referred to as an extrusion plastometer

(Tinius Olsen, Willow Grove, PA), was used at PFS. To use this technology,

approximately 4–5 grams of polymer is placed into the loading cell. The material

is heated to 190 EC (374 EF) for 6 minutes. A 2.16 kg (4.75 lb) load is also applied

during the melt, inducing a pressure of 3.0 bar (3032.148 g/cm2

). The molten

polymer is allowed to flow through an orifice for 1 minute, and the extruded plastic

is then weighed. The melt flow index (MFI) is the output rate in g/10 minutes from

a standard die of 2.1 mm diameter and 8.0 mm length. The MFI is an inverse

measure of melt viscosity. Therefore, the higher the melt flow index, the lower the

average molecular weight.

The PF111 powder has a melt-flow index of 32 g/10 min and a density of 0.934

g/cm3

. In contrast, PF113 has a higher melt-flow index, 500 g/10 min and a lower

density, 0.918 g/cm3

.
USACERL TR 98/14 13

*

spherulitic: crystals arranged in the general form of a sphere. ** sputtering: a technique for depositing metal atoms on plastic or other substrate.

LaserStrobeTM Control Vision

Particle velocities were measured using a LaserStrobeTM Control Vision System

(Idaho Falls, ID). This system uses two nitrogen laser strobes (5 ns pulse width)

with fiber optic cable, to transport and focus 337 nm light to the area of interest.

The high speed electronic shutter (50 ns to 5 μs) is synchronized with the laser

flashes. The laser is also synchronized with the framing of the video sensor (5 ns

at 337 nm) and is fired once for each captured video frame. The image of the

moving particle is effectively frozen in space and—in double exposure—displaced

in flight. Particle velocity is calculated from the laser synchronization time and

particle displacement [15].

Light Microscopy (LM)

Material specimens were cut with a diamond saw, mounted in cross-section, and

polished using a Buehler automated polishing machine. A two-light microscope

(Unitron-Versamet, Plainview, NY) was used in reflection to observe coating

microstructures. Lamellae (crystalline regions) were best visualized when using

darkfield illumination. This was accomplished by setting the epi-diaphragm lever

to “D,” adjusting the polarizer to 90 degrees, and removing both the analyzer and

the interference prism from the light path.

Coatings were also sectioned using a microtome to obtain optically transmissive

specimens. Polarized optics were used in transmission to detect spherulitic*

features. Anisotropy is present when black and white contrast is visualized.

Polarized optics were obtained by setting the epi-diaphragm lever to “B,” setting the

polarizer to 0E setting the analyzer to 90E and removing the interference prism from

the image path.

Scanning Electron Microscopy (SEM)

Polymer powder, coating surfaces, and coating cross-sections were sputtered** with

gold and analyzed with an ISI-SX-30 scanning electron microscope. The

accelerating voltage was 15 kV (unless otherwise stated in text). Secondary

electron images were used to examine fracture surfaces, and back-scattered electron

images were used to assess pigment agglomeration and other chemistry-dependent

features.
14 USACERL TR 98/14

Surface Profilometry

The surface roughness, Ra, was measured using a Mitutoyo SurfTest III (Japan).

This device uses a mechanical stylus to traverse the specimen surface and measure

the average deviation of the surface from a centerline value. Various scanning

speeds and cutoff values can be selected to examine the surface. General surface

waviness can be assessed using a 6 mm/sec scan rate and a 2.5 cutoff value.

Roughness within the surface undulations can be assessed using a 2 mm/sec scan

rate in conjunction with a 0.08 cutoff value.

Differential Thermal Analysis (DTA)

Polymer powder and coatings were analyzed with a Stanton Redcroft STA 780.

Approximately 10–20 mg of material was placed in a platinum crucible and heated

from 25 to 600 EC (77 to 1112 EF) at a heating rate of 5 EC /min. The constant

heating rate is applied to both the experimental sample and a reference sample.

Because energy input remains constant, differences in temperature between the two

samples are measured as a function of temperature or time as the sample is heated.

The temperature peaks are related to enthalpic reactions due to both physical and

chemical changes to the polymer structure.

Thermogravimetric Analysis (TGA)

Thermogravimetric analysis was conducted simultaneously with DTA on the

Stanton Redcroft STA 780. This technique records the weight of a sample as a

function of time or temperature. As previously discussed, the heating rate in this

study was 5 EC/min. Data pertaining to polymer volatization and decomposition

were obtained.

Differential Scanning Calorimetry (DSC)

Differential scanning calorimetry was performed on a Dupont 2100 V4.OB unit.

This technique is based on heating or cooling an experimental sample and a

reference sample at a predetermined rate. Unlike DTA, the temperature of both the

experimental sample and the reference sample is kept constant while the heat flux

is measured. The heat flux is directly related to the enthalpy, so, heats of reaction

can be quantified. Approximately 8–10 mg samples of polymer were heated from

-30 EC to 150 EC at 10 EC per minute. DSC was used to determine the

temperatures at which the ploymer transitions from a hard, brittle material to an

elastoner (glass transition temperature) and from an elastoner to a liquid (melt

temperature).
USACERL TR 98/14 15

QUV Weathering Chamber

The QUV accelerated weathering tester is a light- and water-exposure testing

apparatus that simulates the damaging forces of ultraviolet (UV, from sunlight) and

rain. Coatings were sprayed in triplicate onto 4 in. X 6 in. (approximately 10 X 15

cm) abrasive-blasted steel of 1/8 in. (3.2 mm) thickness and subjected to accelerated

QUV weathering tests according to ASTM G 53-88. A carbide-tipped scribe was

used to scrape an “X” on the coating surface. Samples were exposed to cyclic

conditions of ultraviolet light, 280-315 nm (UV-B) at 60 EC (140 EF) and to

condensation at 50 EC (122 EF) every 4 hours for a total of 1250 hours. Samples

were rotated once a week to ensure uniform exposure. Ultraviolet lamps were

rotated every 500 hours in accordance with manufacturers specifications.

Tensile Testing

Techniques were developed to produce freestanding polymer coatings. The first

method involves painting a water-soluble debonding layer (Metco De-Bond,

Westbury, NY) on the surface of a steel plate and allowing the layer to dry for 24

hours. Polymer powder was then sprayed directly onto this debonding layer;

however, the propane flow rate was decreased during the first few passes to prevent

vaporization of the debonding layer. The propane flow was then increased to the

desired rate and coatings were deposited to the appropriate thickness. Once the

polymer cooled, an exposed edge was immersed in water, allowing the coating to be

easily removed from the substrate. Although freestanding polymer components can

be prepared in this fashion, high application temperatures degrade the debonding

agent, and this produces porosity in the polymer coating.

The second method of preparing free-forms entails spraying the molten polymer

directly onto a TeflonTM-coated skillet. Since the fluoropolymer has such a low

surface energy, the polymer coating does not adhere to the skillet surface allowing

easy removal. This method has significant advantages over the water-soluble layer

method and was thus used for all testing of polymer mechanical properties.

To minimize surface flaws that may arise from cutting, a die-stamp was used to

produce reproducible ASTM D632 Type IV “dog-bone” specimens. The tensile test

specimens were also polished to ensure uniform sample thickness. Tensile testing

was conducted on a system manufactured by Applied Test Systems (ATS). A 1000

lb load cell measured the load to maintain a constant displacement rate. Data were

collected at 80 Hz using a data acquisition package. Load and displacement values

were converted to engineering stress and strain using a strain rate of 2 inches per

minute (5 cm/min). The effects of pigment, melt-flow index, and cooling rate were
16 USACERL TR 98/14

also assessed. A minimum of eight samples were tested for each condition to

establish a standard deviation.

Peel Adhesion Testing Rig

In order to assess the polymer adhesion to steel, a 90 degree peel test rig was

constructed according to ASTM D 3167 specifications. The geometry consists of two

parallel plates held one in. apart by three studs (Figure 3). Two 1-in. diameter

mandrels are set in needle roller bearings to allow the substrate to traverse as the

load is applied. This construction permits the jig to align itself as well as allow the

sample to climb during the experiment.

Steel substrates measuring 6 in. in length, 1 in. in width, and 1/8 in. thickness were

used in this study. Self-adhering aluminum tape was initially fastened to one end

of the steel substrate. A polymer coating of approximately 625 micrometers (25

mils) was then deposited over the tape/steel assembly. The extended tab was

fastened to the grips of an ATS tensile testing machine and subsequently pulled 90

degrees from the steel substrate. This procedure allows the crack to propagate

along the polymer/steel interface at the peel speed, specified as 152 mm/min (6

in./min). Figure 4 illustrates the crack direction as the plastic adherate is peeled

from an abrasive-blasted steel substrate. The force required to propagate the crack

was recorded on a digital acquisition board at a rate of 8 Hz.
USACERL TR 98/14 17

Figure 1. Schematic of combustion spray torch.
18 USACERL TR 98/14

Figure 2. The effect of gas mixture upon (a) flame temperature and (b) flame length.
USACERL TR 98/14 19

Figure 3. Construction of peel test jig.

Figure 4. Peeled polymer coating from abrasive-blasted

substrate.
20 USACERL TR 98/14

Effect '

EY%

n%

&

EY&

n&

(1)

3 Experimental Results

Flame Spray Processing of Virgin Polymers

A design of experiments approach was employed to study the coalescence of

combustion-sprayed ethylene-methacrylic acid (EMAA) copolymer [16]. The EMAA

was obtained from Plastic FlameCoat Systems as PF113W (Figure 5). This powder

was dry compounded with TiO2

 pigment as well as with proprietary thermal and

UV light stabilizers. PF113 has a melt index of 500 g/10 min, corresponding to the

lowest average molecular weight copolymer in the EMAA series. This polymer was

chosen on the basis of its sensitivity to temperature, which allowed precise

processing/structure evaluations to be made.

Thermogravimetric analysis and differential thermal analysis were conducted

simultaneously on the PF113 powder. The first endothermic peak began at 72 EC

and ended at 96 EC, characteristic of the polymer melting point region. Weight loss

initially occurred at 260 EC losing 5 percent of the weight between 260 and 390 EC.

This may signify stabilizer decomposition since an endothermic peak was observed

on the DTA trace at 260 EC. At 390 EC an endothermic peak on the DTA trace was

observed, representing polymer vaporization. This transition was directly reflected

on the TGA spectrum because 75 percent weight loss occurred from 390 to 470 EC.

The thermal analysis therefore suggests that PF113 can be processed between the

temperatures of 94–260 EC (201–500 EF) without weight loss. The maximum

processing temperature before polymer vaporization is 390 EC (734 EF).

Since a large processing window was verified using DTA/TGA, it was hypothesized

that a range of splat morphologies is likely to exist during spray processing. A 25-1

factorial design matrix was chosen for EMAA copolymer to determine the processing

factors responsible for changes in the coating microstructure. Using a two-level

design, process parameters were varied from a low value (-) to a high value (+), and

the effect upon a particular response (Y) is measured [17]. The “effect” is calculated

by subtracting the average response using the low parameter from the average

response using the high parameter setting, i.e.,
USACERL TR 98/14 21

Yi' $o%$1X1%$2X2%$12X1X2%$3X3%$13X1X3%$23X2X3......%$iNXiXN (2)

$ '

Effect

2

(3)

Xcoded '

X&[(%)%(&)]

[(%)&(&)]

(4)

Interactions between parameters can also be detected and quantified [17]. Influential process parameters can, therefore, be identified on the basis of the induced

effect. The underlying objective of empirical modeling is to relate N coded process

parameters Xi (I=1...N; here: N =5) to M responses Yj (j=1...M; here: M=3) of the

form

where the beta coefficients are regression coefficients calculated from the measured

effect [14], i.e.:

The “Xi

” terms in the polynomial equation are coded process parameters

representing the selected parameter value multiplied by an empirical factor

calculated from the chosen high and low values (+ and -, respectively):

This allows prediction of a particular response on the basis of the selected process

parameters.

The present model entails setting five parameters (powder feed rate, standoff

distance, substrate preheat temperature, propane flow rate, and compressed air

flow rate), at two levels (+ and -), measuring three responses (surface roughness,

coating temperature, and splat elongation ratio), and calculating the effects due to

each parameter and parameter interaction. A design of experiments software

package, Design-Ease (Statease Inc., Minneapolis, MN), was used to generate the

half-factorial matrix, calculate the single parameter and interaction effects, and

provide the analysis of variance (ANOVA). Three additional sets of parameters

using mean high / low values were also used to account for nonlinearity as well as

to determine the repeatability of the deposition process, giving a total of 19

experiments. A number of fixed and variable parameters were selected on the basis

of thermal spray experience (Table 1). A six-axis articulated robot was used to

ensure that the traverse speed, standoff-distance and spray step distance remained

constant during spraying. Although a traverse speed of 25 cm/sec is faster than one
22 USACERL TR 98/14

Ytemperature ' 118.3%11.5X1coded%5.8X2coded%33.1X3coded%12.5X4coded&6.1X2codedX5coded (5)

would normally use in practice, it was used to ensure minimal flame contact with

the coating.

Upon completion of each experiment in the matrix, a remote infrared pyrometer

measured the surface temperature of the sprayed deposit (noted as response 1).

The surface roughness, Ra

, was measured as previously described in section 3.7

(noted as response 2). The microstructure was visualized with light microscopy.

Buehler image analysis software (Buehler, Lake Bluff, OH) was used to trace the

individual polymer lamellae, and the results were then digitized for subsequent

measurements. Thirty splats per sample were analyzed to determine the average

elongation ratio of the splat (noted as response 3).

The components of the 25-1 factorial matrix as well as the measured responses for

each set of experimental conditions are listed in Table 2. Using a traverse speed of

25 cm/sec, coating surface temperatures ranged from 60–180 EC, surface roughness

ranged from 50 micrometers to 3 micrometers, and splat elongation ratio ranged

from 1.59 to 7.33. Table 3 summarizes the parameters and parameter

combinations inducing the greatest effects on the responses. The powder feed rate

had the greatest impact on coating microstructure and macrostructure (splat

morphology and surface roughness, respectively). Substrate preheating had the

greatest effect on the coating surface temperature.

Coating Temperature

The temperature of the deposited thermoplastic coating depends upon the substrate

preheat temperature, X3

, the propane flow rate, X4, and the powder feed rate, X2.

Standoff distance, X2

, and the interaction between standoff distance with

compressed air, X25, are also included in the model. An increase of substrate

temperature from ambient to 87 EC increased the coating temperature from 85 to

151 EC (effect = 66 EC) when averaged over all parameter combinations in the

matrix. Increasing the propane flow rate from its low flow setting to the high flow

setting increases the coating temperature from 106 to 131 EC (effect = 25 EC). Since

the length of the flame increases with the propane flow rate, higher particle

temperatures result due to greater residence times within the flame. The average

coating temperature increased from 107 to 130 EC (effect = 23 EC) when the powder

injection rate increased from 70 g/min to 140 g/min. The polynomial equation

which empirically predicts the coating temperature (in EC) to within 5 percent is:
USACERL TR 98/14 23

X1coded '

X1&105

35

(g/min) (6)

X2coded '

X2&37.5

12.5

(cm) (7)

X3coded '

X3&54

33

(

oC) (8)

X4coded '

X4&10.9

4.3

(l/min) (9)

X5coded '

X5&160

40

(l/min) (10)

where the coded Xi terms are as shown below and are applicable to each of the

empirical equations representing coating temperature, surface roughness, and splat

elongation ratio.

The accuracy of this equation was assessed by using the process parameters from

the design of experiments study (Table 2) and comparing the measured response to

the predicted response. The 5 percent error represents the average residual from

the experimental matrix.

Surface Roughness

The roughness of EMAA coatings depends upon the powder feed rate X1

, substrate

temperature X3

, substrate temperature/propane flow rate interaction X34 and

powder feed rate/standoff distance interaction X12. The propane flow rate X4

 was

also included in the model for increased accuracy. An ln (Y2

) transformation was

used to model the surface roughness due to the large variation in roughness values.

The surface roughness decreased from 32 micrometers to 12 micrometers (effect =

20 micrometers) when the powder feed rate was increased from 70 g/min to 140

g/min. The surface roughness also decreased from 27 micrometers to 16

micrometers (effect = 11 micrometers) as the substrate temperature was increased

from ambient to 87 EC.
24 USACERL TR 98/14

1nYroughness ' 2.81&0.49X1coded&0.23X1codedX2coded&0.31X3coded&0.21X4coded&0.22X3codedX4coded (11)

YE.R.' 3.99%0.75X1coded %0.59X2coded %0.51X1codedX2coded %0.58X1codedX5coded &0.37X2codedX5coded %0.41X3coded %0.34X3codedX4coded %0.35X4codedX5coded(12)

Since both parameters are involved in significant interaction effects, it is essential

to include such interactions in the empirical model. The surface roughness

remained constant at low propane flow rates, even when the substrate temperature

increased, but, the roughness decreased with increasing substrate temperatures at

the high propane flow rate (effect = 9 micrometers). The second noteworthy interaction involves standoff distance and powder feed rate. The surface roughness

decreased with increasing powder feed rate at both 25 and 50 cm. At lower feed

rates, the roughness is lower for shorter standoff distances, but at higher powder

feed rates, the roughness is lower at longer standoff distances (effect = 7

micrometers). The empirical equation used to estimate coating surface roughness

(in terms of Ra in micrometers) is as follows:

Upon analysis of the residuals, Equation 11 was found to be accurate to within 10

percent of the measured surface roughnesses.

Microstructure: Splat Elongation Ratio

The splat elongation ratio depends on powder feed rate X1

, standoff distance X2,

powder feed rate interacting with standoff distance X12, and substrate preheat

temperature X3

. A number of secondary effects X15, X25, X34, and X45—are also

included. The largest effect was observed when the powder feed rate was increased

to 140 g/min and thus increased the average elongation ratio from 3.3 to 4.8 (effect

= 1.5). The average elongation ratio increased from 3.4 to 4.6 (effect = 1.2) when

the standoff distance is increased from 25 cm to 50 cm. The elongation ratio also

increases from 3.6 to 4.4 (effect = 0.8) when the substrate is preheated to 87 EC.

The powder feed rate was also found to interact with standoff distance. The splat

elongation ratio does not change with standoff distance at the low powder feed rate

of 70 g/min. However, at the higher feed rate of 140 g/min, the average elongation

ratio increased by approximately 2.2 when the standoff distance increased from 25

cm to 50 cm. The empirical equation used to estimate the splat elongation ratio (in

nondimensional units) from knowledge of the chosen process parameters is:

Again, as for the surface roughness equation (Equation 11), using the parameters

from the experimental matrix and comparing the residuals, Equation 12 was found

to be accurate to within 10 percent of the measured elongation ratio.
USACERL TR 98/14 25

Differences in the melting and/or deformation behavior of EMAA copolymer can be

discerned by examining the microstructures in cross-section. Figure 6 illustrates

poor coating coalescence due to insufficient polymer melt-flow, as evident by

spherical splats (line 7, Table 2). The degree of melting and coalescence in Figure

7 (line 16, Table 2) is higher, as evident by thin lamellae and a deformed impact

geometry. The impinging molten polymer will flow around the underlying topography, thereby, increasing the density of splats.

The elongation ratio distributions for the poorly coalesced coating and highly

coalesced coating are shown in Figure 8. The average elongation ratio for the

microstructure in Figure 6 was 1.6 with a standard deviation of 0.4. Ninety percent

of the coating contains particles with elongation ratios less than 2. In contrast, the

average elongation ratio for the coating in Figure 7 was 7.3 with a standard

deviation of 3.6. This coating, which corresponds to the lowest of measured surface

roughness values, reveals that less then 2 percent of the deposit contains particles

with elongation ratios lower than 2. It is common for poorly melted particles to

exist in the coating because particles do not all follow the same trajectory through

the flame, but this analysis quantifies the microstructural variation in thermal

sprayed EMAA coatings.

Response Correlations

A general trend was observed between the average elongation ratio and surface

roughness for each specimen in the matrix. Figure 9 illustrates that a greater

surface roughness correlates to lower splat elongation ratios (and, thus, poor

coalescence). From the design of experiments study, two factors seem responsible

for the poor melt flow characteristics. One factor is the low thermal input to the

polymer due to its short residence time within the flame. This produces insufficient

particle melting and results in low elongation ratios. On the coating surface, low

splat elongation ratios correspond to an agglomeration of fused polymer spheres.

Areas of protruding polymer particulates yield a higher surface roughness, which

defines a poorly coalesced coating. The second possible factor relates to the rate of

polymer injection into the combustion zone. A low powder feed rate correlates to a

lower amount deposited polymer per unit area. Less molten material can conform

to the underlying topography. Discontinuous surface coverage will result in a

higher surface roughness unless a slower traverse speed is used during application.

The temperature dependence upon surface roughness is shown in Figure 10. A high

surface roughness was measured for coating temperatures lower than 80 EC. The

surface roughness decreased by a factor of approximately 2 at temperatures greater

than 90 EC. This temperature threshold between high to low surface roughness
26 USACERL TR 98/14

correlates with the melting point of the copolymer. Data obtained from differential

scanning calorimetry reveals that the primary endothermic melting point begins at

78 EC and peaks at 96 EC. However, the elongation ratio measurements (Figure 11)

were not as sensitive to temperature as the surface roughness measurements. The

elongation ratio only correlated to the coating temperature at the extreme temperature limits. The temperature region of 95-140 EC may be sufficient to reduce the

surface roughness, but temperatures within this range are not sufficiently high to

elongate the polymer splats. At 180 EC a noticeable increase in splat elongation

ratio occurred. It also should be noted that the splat elongation ratio is a function

of the original diameter of the polymer particulate. Larger polymer particles will

remain more spherical in shape than smaller particles. Less energy is required to

melt the fine particles, so smaller particles are more susceptible to deformation at

the site of impact than the larger polymer particulates. Examining the correlations

between the responses, a well-coalesced EMAA coating will result from an application temperature greater than 125 EC, and will exhibit a surface roughness less

than 15 micrometers and a splat elongation ratio of at least 4.5.

Particle Velocities

A LaserStrobeTM

 Control Vision system was used to determine whether particle

velocities were dependent upon the process parameters. The experimental technique was the same as previously described. Particle velocities (50 particles per

data point) were measured at several standoff distances for both a high flame setting ting (high propane, high compressed air) and for a low flame setting (low propane, low compressed air). At a distance of 25 cm from the PF200 nozzle, average

particle velocities are 14 m/sec and 10 m/sec for the high and low gas settings,

respectively. The standard deviation was 5 m/sec, indicating that increasing gas

flow rates do not appreciably affect the particle velocity. In addition, particle velocities at 50 cm from the nozzle are only slightly lower than at 25 cm, irrespective

of gas settings. Variances in surface roughness and elongation ratio did not arise

from differences in particle velocities since the velocities were found to be independent of the gas flow rates and standoff distance. The particle residence time varies

from 3 ms for a 5 cm length flame to 20 ms for a 25 cm length flame. The flame

stoichiometry did not appreciably affect the fusion of the polymer particulates.

Discussion of EMAA Thermal Spray Tests

The microstructure of a thermal sprayed polymer deposit is process-controlled. The

substrate preheat temperature showed the largest effect on coating temperature.

An empirical equation was developed to estimate coating temperature to within a

5 percent margin of error. The powder injection rate showed the largest effect on
USACERL TR 98/14 27

*

comminution: reduction of a material to a powder.

coating surface roughness and the splat elongation ratio. Polynomial equations

were experimentally determined for surface roughness and splat elongation ratio,

and were accurate to within 10 percent.

Center-line particle velocities did not change with the gas flow rates or standoff

distances used in this study (approximately 15 m/sec ), but differences in splat

elongation ratio were observed. The particle residence time varied from 3 ms for a

5 cm length flame to 13 ms for a 20 cm length flame. The flame stoichiometry did

not appreciably affect the fusion of the polymer particulates, so, for a fixed traverse

speed, the particle residence time within the flame as well as the number of

particles arriving at the substrate surface determines the extent of EMAA coating

coalescence.

Based upon the correlations between the responses, a well-coalesced EMAA coating

(melt flow index of 500 g/10 min) will have a temperature greater than 125 EC (257

EF), a surface roughness less than 15 micrometers, and a splat elongation ratio of

at least 4.5. The coating temperature, surface roughness and microstructure can

be estimated on the basis of the selected process parameters.

Flame Spray Processing of Recycled Polymer Blends

Processing

Post-consumer commingled polymer (PCCP) was received from Obex, Inc.

(Stamford, CT) in the form of 3/8" flake. PCCP consists principally of different

polyethylene densities and molecular weights, plus smaller concentrations of

polypropylene, polystyrene, polyvinylchloride and polyethylene terepthalate. The

polymer chips were washed to remove food debris and residue. Aluminum, glass,

cardboard, and paper were extracted from the polymer blend. A cryogenic milling

machine using liquid nitrogen coolant was employed to lower the temperature of the

polymer to facilitate communition*

. The ground polymer was sieved to -50 mesh,

corresponding to an average particle size of 177 micrometers (Figure 12).

A combustion torch marketed by Plastic FlameCoat Systems as PF200 was used

with the addition of a fluidized bed powder delivery system. Polymer coatings were

sprayed to a thickness of approximately 800 micrometers (30 mils) at a temperature

of 218-232 EC (425-450 EF).
28 USACERL TR 98/14

Microstructural Analysis

The cross-section of the as-sprayed PCCP coating is shown in Figure 13. Contrast

is observed in the cross-section due to the various polymer chemistries, sizes, and

pigments originally used in the processing. Polymer particles impinging upon the

substrate (preheated to 77 EC [170 EF]) are either semi-molten or fully molten.

Upon impact, the polymer droplet (splat) will spread, deform, and coalesce to the

substrate or underlying topography. The lower melting point polymers such as low

density polyethylene will have a lower melt-viscosity than higher molecular weight

formulations, and thus will flow more easily. Spherical splats are also observed in

the microstructure representing insufficient particle melting.

Due to the large differences in surface tension between different polymeric species

in the PCCP blend, limited diffusion can occur between incompatible splats. Recent

studies have also shown the need for proper melt-compatibilization to improve the

cohesive strength between immiscible polymers [18, 19, 20]. Table 4 lists the

critical surface tensions, solubilities, and thermal properties for the polymers

making up PCCP. The polymer combination with the least compatibility is

polyethylene with polyethylene terphthalate.

The surface of the as-sprayed PCCP coating is shown in Figure 14. Thin

macroscopic cracks, of lengths up to 800 micrometers, propagated along consecutive

splat boundaries. Interfacial boundaries between incompatible polymer species, in

conjunction with differences in thermal expansion behavior is most likely

responsible for inducing the cracks between neighboring splats during cooling.

Accelerated Degradation

The QUV accelerated weathering tester simulates the damaging forces of sunlight

and rain. Coatings were exposed to 4 hours UV-B light at 60 EC and 4 hours condensation at 50 EC for a total of 1250 hours. Deterioration to the surface includes

cracking, chalking, and photo-bleaching. The extent of these distresses depends

upon the volume fraction of EMAA blended with the PCCP. Figure 15 shows

weathered surfaces of 100 percent PCCP, 50 percent PCCP, 25 percent PCCP, and

100 percent EMAA. Figure 16 depicts the dominant crack propagation paths.

For the case of 100 percent PCCP, both the unmelted and melted PCCP particles

have poor interfacial adhesion to the surrounding matrix, resulting in interfacial

debonding. Cracks were observed with lengths of up to 1600 micrometers. These

longer cracks seem to propagate tangentially from the particle, in the direction of

the interfacial disbondment. Cracks shorter in length also initiated radially from
USACERL TR 98/14 29

the particle/matrix interface. The 75 percent PCCP coating appeared similar to the

100 percent commingled case, but the interfacial zone between particle and

surrounding polymer was slightly narrower. Radial cracks from one particle or

splat intersected other radial and tangential cracks from neighboring particles. The

PCCP particles are distributed farther apart in the 50/50 mixture, thus decreasing

the number of crack intersections. Unmelted particles are discernible, but they are

clearly embedded within the surrounding matrix. Radial cracks from one particle

often propagated the distance to surrounding incompatible polymer splats. At a 25

percent PCCP volume fraction, only radial cracks were observed to propagate from

the PCCP/EMAA interface, and inter-particle cracking was not evident. Cracks

were not observed for the pure EMAA coating, but photo-bleaching was evident,

resulting in a loss of gloss.

The cracking was measured to quantify the extent of degradation to the PCCP

blends. A pseudo crack density index was determined using a five-line intercept

method (Figure 17). The error bars represent the standard deviation from

measuring triplicate samples. The number of cracking events decreased as the

percentage of virgin EMAA copolymer in the blend was increased. The 75 percent

PCCP blend had approximately the same crack density as the pure commingled

polymer. The 25 percent PCCP blend had a crack density index of 0.3 cracks/mm

(3 cracks/cm). Crack length distributions were also determined for each coating

system using optical microscopy, and the average crack length was plotted in Figure

18. Each sample had a large distribution of cracks, and therefore had a large

standard deviation. However, the overall trend suggests that increasing the volume

percentage of EMAA in the blend decreases the average crack length. These

numbers do not include cracks smaller than 10 micrometers, which were later

observed using scanning electron microscopy.

Adhesion of PCCP and Blends

In order to assess coating adhesion to steel, the standard ASTM C633 protocol was

modified. It was previously found that epoxy does not bond well to thermoplastic

surfaces primarily due to low surface energies. The tensile adhesion test stud was,

therefore, abrasive-blasted to 10 micrometers, heated to 204 EC (400 EF) and then

placed on the coating with a force of 350 g. This allowed the stud to be fusion

bonded to the coating surface. The circumference of the stud was also scribed to

bare metal to eliminate the possibility of shear forces being distributed to adjacent

polymer during the test. The hydraulic Instron used a constant strain rate of

0.0005 in/sec (approximately 0.0013 cm/sec) and the force required to remove the

coating was recorded.
30 USACERL TR 98/14

Unblended PCCP and PCCP blends with EMAA copolymer were sprayed to the

same thickness (30 mils) onto steel substrates with a 10 mil anchor profile. The

tensile adhesion strength was measured before and after weathering (Figure 19).

The data points reflect an average strength derived from five measurements.

Failure mechanisms changed from 100 percent adhesive (for the 100 percent PCCP

coating) to 100 percent cohesive failure (for 100 percent virgin EMAA) as the

percentage of EMAA increased in the commingled polymer coatings. The

weathering-induced cracking did not significantly affect the adhesive properties of

the applied coatings. Rust was also evident on the steel surface from which the 100

percent PCCP coating was removed, signifying water penetration. It can be

concluded from the data that the adhesion to steel increases as the volume

percentage of EMAA in the PCCP blend increases.

Discussion of PCCP Flame Spray Tests

Microstructures of as-sprayed PCCP coatings show evidence of cracking.

Accelerated weathering induced further cracking, photo-bleaching, and a higher

surface roughness. The PCCP powder was blended with ethylene methacrylic acid

copolymer in ratios of 25 percent, 50 percent and 75 percent by volume.

Introduction of the EMAA was shown to decrease the weathering-induced crack

density as well as reduce the average crack length. Cracking was not evident in the

100 percent EMAA coatings. Blending EMAA copolymer with PCCP also increased

the adhesion to steel from 250 psi for 100 percent PCCP to 1100 psi for 25 percent

PCCP. Accelerated weathering did not appreciably affect coating adhesion. Using

unmodified post-consumer commingled polymer in volume fractions less than 25

percent may promote significant materials and waste disposal cost savings, but not

without a sacrifice in coating cohesion.

Mechanical Properties of Polymer Materials and Blends

EMAA Copolymer

The purpose of this study was to determine the tensile properties of spray-formed

coatings and determine whether pigment or cooling rate affects these properties.

Two EMAA copolymer powders were chosen, having melt flow indices of 500 g/10

min (PF113) and 32 g/10 min (PF111). As previously discussed, coatings were

sprayed onto a Teflon®

 coated skillet that was preheated to 66 EC (150 EF). Once

cool, the deposit is easily removed and stamped to the ASTM D 632 type IV

dog-bone geometry. Table 5 summarizes the average secant modulus (Young’s
USACERL TR 98/14 31

modulus), yield strength, yield strain, tensile strength, elongation to break, and

toughness values obtained for each EMAA coating system.

When PF111-clear was deposited at 160 EC (320 EF), the splats had not completely

coalesced, resulting in a rough surface. The tensile strength, elongation to break,

and toughness were considerably lower than coatings deposited at 216 EC (420 EF),

however, the modulus and yield strength were not appreciably different. During

testing, a neck forms at the yield point and then propagates over the gauge length.

Thus, PF111 coatings at 160 EC cannot plastically deform due to lower splat cohesion in comparison to coatings sprayed at 216 EC. A smooth surface was apparent

without visible porosity for PF111 at 216 EC. In addition, coatings elongated to

approximately 800 percent elongation before failure. Since this polymer strain

hardens (Figure 20) the tensile strength increases with increasing strain. The

presence of pigment increased the modulus and yield strength only slightly, but the

pigment also reduced the elongation to break decreasing both the tensile strength

and the toughness. When unpigmented PF111 was sprayed at 216 EC and quickly

quenched into water at 15 EC, a highly transparent coating was produced. In

comparison to normal cooling (Figure 21) the quenched sample had a lower

modulus, lower yield strength, higher tensile strength, higher elongation to break,

and higher toughness. These data suggest that the crystallinity and the density

had decreased, resulting in a more open structure, thus decreasing the modulus.

The elongation may have increased due to fewer entanglements, resulting in

greater chain orientation during tensile testing. The tensile strength increases with

increasing elongation to break due to strain hardening. When PF111 unpigmented

coatings were sprayed at 271 EC (520 EF), the coatings appeared yellowish in color

due to an increase in oxidation products. Spraying at this temperature also induced

visible porosity in the coatings resulting in 100-500 micrometer bubbles. These

bubbles effectively reduce the cross-sectional area, explaining the decreased

modulus, yield strength and tensile strength. Interestingly enough, the bubbles

deformed in the direction parallel to the applied stress and did not reduce the

elongation to break; thus, only a minor sacrifice in toughness was observed.

The copolymer PF113 was also sprayed at various processing conditions. PF113,

which has a higher melt flow index than PF111, results in a coating that has a

lower modulus, lower yield strength, lower tensile strength, lower elongation to

break, and lower toughness than PF111 coatings. The shape of the stress-strain

curve is quite different as well, due to the lack of strain hardening. Once the yield

point is reached, the tensile strength essentially remains constant during necking.

However, when PF113 was water-quenched at 160 EC, not only did the elongation

to break increase, but the molecular architecture changed due to the observed

strain hardening effect (Figure 22). Since PF113 has a higher melt-flow index, and
32 USACERL TR 98/14

thus a lower molecular weight, coatings could be deposited at lower temperatures.

Mechanical properties improved as coating temperatures increased from 104 to 160

to 216 EC. Once again, the modulus and yield point did not differ at these

temperatures, implying that only plastic deformation is dependent upon the

microstructure. Pigment slightly influenced the modulus and strength, but did

decrease the elongation to break. Figure 23 summarizes the properties of various

PF113 coatings deposited at 160 EC.

Recycled PCCP Blends

Post-consumer commingled plastic coatings and blends were sprayed with the

Plastic FlameCoat 400 torch with gravity feed. The PCCP powder was previously

melt-compounded and screened to a -50 mesh. EMAA copolymer powder (PF111-

pigmented) was mechanically blended in a salt-and-pepper mixture in increments

of 10 percent by volume. As the composition varied from 0 to 100 percent PCCP,

coatings were deposited at increasing temperatures, ranging from 204 EC (400 EF)

at 0 percent PCCP to 232 EC (450 EF) at 100 percent PCCP loading. Table 6

outlines the mechanical properties of such blends and Figure 24 depicts the

stress-strain curves for each mixture. Since the matrix phase varied from 0 to 100

percent, the entire spectrum is represented, i.e., elastomeric particles dispersed in

a brittle matrix and reinforcing particles in an elastomeric matrix. Each mechanical property was then plotted as a function of EMAA loading (Figure 25). Although

100 percent PCCP coatings have a higher modulus and greater tensile strength

than 100 percent EMAA coatings, PCCP coatings do not have much ductility. A 10

percent addition of PCCP decreased the elongation to break from 637 percent to 114

percent. This effect reduces coating toughness from 8868 lb/inch to 1354 lb/inch.

The secant modulus increased linearly with increasing PCCP content. In addition,

tensile strengths increased with increasing PCCP contents up to 70–80 percent

PCCP. The tensile and yield strengths did not change with further PCCP loadings.

The lowest tensile strength corresponded to a 10 percent addition of PCCP. This

10 percent inclusion had the additional affect of preventing the EMAA copolymer

to strain harden during extension. This also explains why the yield strength data

show a different trend than the tensile strength data at low PCCP loadings.

Discussion of Mechanical Properties Findings

The mechanical properties of combustion-sprayed polymers depend upon the

processing parameters. The processing parameters determine coating temperatures

which, in turn, affect the polymer microstructure. If the deposition temperature is

too low (104 EC for PF113 and 160 EC for PF111), coatings have low strengths, low

elongation ratios, and low toughness values. Increasing the coating temperature
USACERL TR 98/14 33

allows the particles to fully coalesce, resulting in maximized strength and

toughness. However, at 271 EC, PF111 had visible porosity, which decreased the

strength and modulus of elasticity. Pigment acts as reinforcement in the sense that

the modulus increased. However, the elongation to break decreased, reducing

overall coating toughness. It is believed that water quenching decreases the

crystallinity, resulting in a lower elastic modulus and yield strength, but higher

elongation to break.

Coating toughness was decreased by blending PCCP with EMAA copolymer. The

decrease in toughness and elongation to break was accompanied by an increase in

strength and elastic modulus. Since PCCP contains approximately 80–90 percent

polyethylene and EMAA is a functionalized polyethylene, some compatibility is

expected to exist.

EMAA Adhesion to Steel

Peel Strength

In order to assess the polymer adhesion to steel, a 90 degree peel test was constructed according to ASTM D 3167 specifications, as previously discussed. Such

tests allow the determination of the localized bond strength at the coating

/substrate interface [21-23]. The measured peel force is the sum of the force

required to mechanically bend the coating and the force to break secondary bonds

at the interface. The force required to propagate a crack of width equal to the

specimen width is recorded on a digital data acquisition board at a rate of 8

measurements per second.

The significance of molecular weight on adhesion was investigated by comparing

two formulations of EMAA—PF111 and PF113—having melt flow indices of 32 and

500 g/10 min. The copolymers were obtained from Plastic FlameCoat, Inc. as

PF111 and PF113, respectively. The virgin EMAA powders were flame sprayed

with a Plastic FlameCoat 200 gun 7, using propane and compressed air as the fuel

gasses. The spray conditions were chosen to fully coalesce the particles well within

pyrolysis limits as determined from the previous studies. The influence of substrate

topography on peel strength was ascertained by spraying steel substrates under the

following surface preparation conditions: polished, abrasive-blasted to an Ra of 10

micrometers, and the as-received condition.

Figure 26 summarizes the average peel strengths for PF113 and PF111 as a

function of steel topography [24]. The peel strength values reflect the averaged
34 USACERL TR 98/14

force required to separate the two surfaces, which includes the energy dissipated by

plastic deformation. Higher peel strengths were observed for the higher molecular

weight EMAA copolymer for each steel surface profile. The peel adhesion strength

increased fourfold for PF113 upon abrasive-blasting the as-received steel and

increased by a factor of two for PF111. Peel strengths also increased (for both

formulations) by approximately 1000 J/m2

 upon polishing the as-received steel.

Polished Steel

The peeling fracture mechanisms were dependent upon the substrate topography.

In a series of experiments designed to determine the effect of topography on the

performance of the coatings, EMAA copolymer was applied to polished steel,

abraisive-blasted steel, and on steel as received from the supplier. The EMAA

copolymer coatings on polished steel show unstable crack propagation, as evident

by the sawtooth load-displacement curves in Figure 27. Such behavior was

observed for both PF111 and PF113 materials, but a greater force was required to

propagate and bring the crack to rest for PF111. The peaks on the load-displacement plot correspond to crack initiation and the valleys correspond to crack arrest.

This implies that the crack accelerates faster than the strain rate, 6 mm/sec, and

the strain energy release rate, G, is larger than the interfacial fracture toughness,

'

5

. The crack continues to propagate until the strain energy release rate is

balanced by the increase in fracture surface area.

The fracture mechanism is referred to as “slip-stick” failure. Polymer was extracted

from the coating creating a series of parallel band markings on the surface of the

polished steel. Examination of the polymer surface from which the polymer was

extracted also reveals band markings (Figure 28). However, using higher magnification, the markings on the polymer surface actually represent a series of crack

networks. Not only did the position of the band markings coincide between the

polymer coatings and the respective steel substrates, but their position also

corresponds to the minima on the load-displacement plot. Polymer material was

only found within the discrete band locations on the steel surface, signifying that

splat decohesion did not take place during crack propagation. The distance between

the markings for both PF111 and PF113 increased linearly from 1 mm to

approximately 5 mm during the test. This was observed by measuring the distance

between the markings on the steel surface and was verified by measuring the

distance between peak minima. This suggests that a greater amount of energy is

absorbed and subsequently released as kinetic energy as the displacement

increases.
USACERL TR 98/14 35

Abrasive-Blasted Steel

Polymer strips peeled from an abrasive-blasted surface with a surface roughness,

Ra

, of 9 micrometers failed in a different manner. As evident in Figure 29, the

output does not show the same degree of load deflections that occurred upon peeling

from polished steel. In addition, polymer band markings were not observed on the

surface of the roughened steel nor on the coating. Examination of the polymerside

of the interface (post-peeling) revealed a highly deformed surface. In fact, the

higher molecular weight polymer coating (PF111) showed a greater amount of

deformation than PF113. In addition, a greater peel force was recorded for PF111.

Holes on the surfaces of both coatings were evident, surrounded by a plastically

deformed rim (Figure 30). These holes signify splat decohesion from the matrix

since the holes are about the same size as a splat (approximately 100 micrometers).

It appears that the matrix attempted to constrain the splat but failed. Polymer

fibrils were often observed near the perimeter of such holes. Examination of the

abrasive-blasted steel surface revealed that polymer material was clearly anchored

within the undulations (see Figure 31). The polymer within the steel topography

was also elongated to a fibrous geometry. This suggests that the polymer was

sufficiently anchored to the steel to allow a filament to be formed and ultimately

withdrawing the particle from the coating. A fibril or filament near a hole on the

polymer surface may actually be the remains of a splat after extension and

subsequent rupture. A distribution of polymer was found on the abrasive-blasted

steel surfaces for both PF111 and PF113 coatings.

It is apparent that a different fracture mechanism is taking place than peeling on

a polished surface. This failure can therefore be characterized as exhibiting greater

crack stability, although minor load deflections were visible on the load-displacement plots. The energy available for crack propagation is utilized as fracture

surface energy as well as the energy required to extract the splat from the coating.

The minor load deflections may possibly be attributed to the extension of the

polymer splat from the coating to the steel surface. A load drop would then

correspond to the final stage of splat decohesion. Slip-stick failure could not take

place because work was constantly required to remove the mechanically anchored

polymer from the steel topography. This extra work is reflected in the higher peel

force (for both EMAA formulations) to remove the coating from abrasive-blasted

steel as compared to removing a coating from a polished steel surface.

As-Received Steel

The as-received steel specimens contained an undetermined amount of embedded

grease and various amounts of surface oxidation. Failure mechanisms were
36 USACERL TR 98/14

dependent upon the extent of surface contamination. All low molecular weight

EMAA specimens (PF113) peeled in a stable fashion (no load deflections), fracturing

near the interface. Fracture of the higher molecular weight EMAA (PF111) was

dependent upon the location of the surface staining. Regions of the steel containing

embedded grease displayed stable crack propagation. The contamination was

observed on the peeled strip as well. However, areas with less surface staining

displayed failure mechanisms characteristic of slip-stick, with accompanying band

marks on the steel as well as load deflections on the output (Figure 32a). It is not

known what concentration of contamination is required to change the crack front

stability.

Contamination from the as-received steel was found on both PF113 and PF111

polymer coatings (Figure 32b). The peeled polymer surfaces showed significantly

less deformation than coatings peeled from abrasive-blasted surfaces. However,

holes were observed on the peeled surface adjacent to the removed contamination.

Thus, the adhesion to an as-received steel surface was sufficiently strong to remove

some polymer material from the coating. Polymeric splats and filaments were thus

observed on the as-received steel surface. Although the polymer will adhere to

surface oxides, maximum adhesion is achieved by removing surface contamination

and abrasive-blasting the surface to a rough profile. Contaminants and oils will

decrease the interfacial interactions between the polymer and steel surface, thus,

decreasing the interfacial adhesion.

Interfacial Bonding

As previously noted, the experimentally measured peel force is the sum of the forces

required to mechanically bend the coating and those required to break the interfacial bonding. This interfacial adhesion is dependent upon the number of atomic

bonds per unit area multiplied by the atomic bonding energy. In the absence of

primary bonding, Lewis acid-base interactions provide a mechanism for tenacious

interfacial bonding. Fowkes has reported such bonding between various polymers

and substrate materials [25, 26].

One hypothesis can be based upon examination of the polymer chemistry. The

carbonyl moiety, stemming from the carboxyl group, may act as a nucleophile thus

capable of donating electron density (Lewis base) to the surface of steel. The

surface of steel is hydroxylated to some degree and will accept the electron density

since hydrogen is partially positive (Lewis acid). Hydrogen bonding is therefore a

subset of Lewis-acid base reactions. Figure 33 illustrates the proposed secondary

bonding interactions between EMAA and the steel substrate. It should be noted
USACERL TR 98/14 37

Figure 5. Ethylene methacrylic acid copolymer powder.

that a similar model for cohesion is plausible since the carbonyl group of one

polymer chain may attract the hydroxyl group of another chain.

Abrasive-blasting the steel surface increases the surface area and provides

increased mechanical anchoring. The surface energy is also expected to increase

due to bond deformation. A monolayer of oxidation products will form within

minutes of abrasive-blasting and thus contribute to interfacial bonding. However,

porosity may result if polymer is sprayed onto oxide layers greater than a few

monolayers in thickness. Air trapped by oxidation products on steel can diffuse into

the polymer, creating voids of millimeter dimensions. Interfacial porosity provides

a path for crack propagation and thus decreases the interfacial fracture toughness.

On as-received steel surfaces, contamination includes both embedded dirt and

machining oil. Oils are nonpolar and do not provide a site for tenacious bonding.

Embedded dirt and oils act as a lubricant and decrease the interfacial interactions

between the steel and the polymer. This hypothesis is supported by the interfacial

failure mechanism exhibited by a heavily contaminated surface.

Discussion of EMAA Adhesion Findings

The ASTM D 3167 peel test is not suitable for evaluating PCCP blends due to their

lower strength and ductility, but it is quite useful in assessing the adhesion of

flame-sprayed EMAA coatings to various steel surfaces. The copolymer PF111 is

of higher molecular weight and density than PF113 and required a greater load to

propagate an interfacial crack. An abrasive-blasted surface produced the strongest

joint, followed by polished steel and then the as-received steel surface. It is

therefore recommended to abrasive-blast the steel surface for maximum coating

adhesion.
38 USACERL TR 98/14

Fixed Parameters Variable Parameters

Propane tank pressure : 41 kPa Powder feed rate: 70 g/min and 140 g/min

Nitrogen tank pressure: 345 kPa Standoff distance: 25 cm and 50 cm

Compressed air pressure: 827 kPa Substrate preheat temp: 21 EC and 87 EC

Robotic traverse speed: 25 cm/sec Propane flow rate: 6.6 l/min and 15.1 l/min

Spray step distance: 3 cm

Compressed air rate: 120 l/min and

200 l/min

Table 1. Thermal-spray parameters for design of experiments study.

Specimen

Number

Powder

Feed Rate

Standoff

Distance

Substrate

Preheat

Temp

Propane

Flow Rate

Air Flow

Rate

Response

1

Response

2

Response

3

X1

* X2 X3 X4 X5 Yteperature Yroughness YE.R.

(g/min) (cm) EC (l/min) (l/min)

Deposit

Temperatu

re (EC)

Surface

Roughness

(μm)

Splat

Elongation

Ratio

1 70 25 21 6.6 200 60 40 2.06

2 140 25 21 6.6 120 74 22 2.85

3 70 50 21 6.6 120 69 28 4.81

4 140 50 21 6.6 200 96 10 5.13

5 70 25 87 6.6 120 118 20 4.02

6 140 25 87 6.6 200 151 12 4.33

7 70 50 87 6.6 200 121 50 1.59

8 140 50 87 6.6 120 157 10 5.49

9 70 25 21 15.1 120 78 43 2.33

10 140 25 21 15.1 200 108 16 3.85

11 70 50 21 15.1 200 77 48 2.20

12 140 50 21 15.1 120 119 8 5.40

13 70 25 87 15.1 200 160 11 4.21

14 140 25 87 15.1 120 151 11 3.53

15 70 50 87 15.1 120 171 11 4.73

16 140 50 87 15.1 200 182 3 7.33

17 125 38 54 10.9 160 129 13 5.03

18 125 38 54 10.9 160 138 11 4.98

19 125 38 54 10.9 160 131 9 5.98

* The X1

 to X5 and Ytemperature, Yroughness, and YE.R. refer to the variables and responses used in the empirical

modeling within the text.

Table 2. 25-1 design for EMAA copolymer.
USACERL TR 98/14 39

Deposit Temperature (EC)

Surface Roughness

(Ra in micrometers) Splat Elongation Ratio

Substrate Temperature

Effect = 66

Powder Feed Rate

Effect = 20

Powder Feed Rate

Effect =1.5

Propane Flow Rate

Effect = 25

Substrate Temperature

Effect = 11

Standoff Distance

Effect = 1.2

Powder Feed Rate

Effect = 23

Substrate Temperature / 

Propane Flow Rate

Effect = 9

Powder Feed Rate /

Compressed Air Rate

Effect = 1.2

Powder Feed Rate /

Standoff Distance

Effect = 7

Powder Feed Rate /

Standoff Distance

Effect =1.0

Substrate Temperature

Effect = 0.83

* The process parameters are presented in descending order of influence.

Table 3. Statistically significant parameters*.

Figure 6. Cross-section of coating with low splat elongation ratios.
40 USACERL TR 98/14

Figure 7. Cross-section of coating with high splat elongation ratios.

Figure 8. Distribution of elongation ratios for the microstructural extremes.
USACERL TR 98/14 41

Figure 9. Relationship between elongation ratio and surface roughness.

Figure 10. Relationship between surface roughness and deposit temperature.
42 USACERL TR 98/14

Figure 11. Relationship between splat elongation ratio and deposit temperature.

Figure 12. Morphology of post-consumer commingled plastic after grinding.
USACERL TR 98/14 43

Polymer

Glass Transition

Temperature

(EC)

Critical Surface

Tension

(mN/m at 20 EC) 

Solubility

Parameter

103

 [J/m3]

1/2 10-3

Thermal

Conductivity

(cal/s-cm-K)

Polyethylene -80 – -90 32–34 16–17 8–12

Polypropylene -18 34 19 2.8

Polystyrene 100 33 17–19 2.4–3.3

Polyvinylchloride 81 40 19–20 3.5–5.0

Polyethylene

terephthalate 70 43 22

*Ethylene Methacrylic

Acid Copolymer -50 44 5.8

Table 4. Bulk properties of individual polymer constituents.

Figure 13. Cross-section of post-consumer commingled polymer coating.
44 USACERL TR 98/14

Figure 14. Surface of as-sprayed PCCP coatings from dry compounded recycled plastic.
USACERL TR 98/14 45

Figure 15. Surfaces of weathered polymer coatings.
46 USACERL TR 98/14

Figure 16. Cracking mechanisms for recycled post-consumer commingled polymer blends: (a) interfacial

debonding, (b) crack intersection, (c) inter-particle cracking and (d) radial cracking.

Figure 17. Pseudo crack density of exposed coatings.
USACERL TR 98/14 47

Figure 18. Average crack length for exposed coatings.

Figure 19. Tensile adhesion strengths for PCCP coating blends.
48 USACERL TR 98/14

Coating

Secant

Modulus at

5 percent

strain

(psi)

Yield

Strength

(psi)

Tensile

Strength

(psi)

Elongation

to Break

( percent)

Toughness

(lbs/inch)

PF113-pigment

104 EC 8426 ± 592 981 ± 38 1007 ± 43 78 ± 9 698 ± 106

PF113-pigment

160 EC 8081 ± 320 977 ± 23 1136 ± 23 268 ± 21 2789 ± 245

PF113-pigment

216 EC 8137 ± 460 969 ± 30 1115 ± 41 307 ± 42 3161 ± 502

PF113-clear

160 EC 7564 ± 207 970 ± 30 1178 ± 48 372 ± 43 3966 ± 565

PF113-clear

160 EC

H2O quench 5020 ± 327 835 ± 73 1414 ± 113 683 ± 85 7350 ± 1168

PF111-clear

160 EC 8207 ± 456 1007 ± 21 1528 ± 28 407 ± 43 5192 ± 705

PF111-clear

216 EC 8266 ± 465 1074 ± 56 2338 ± 70 783 ± 33 12844 ± 895

PF111-clear

271 EC 6154± 237 936 ± 41 2028 ± 34 806 ± 19 11794 ± 352

PF111-pigment

216 EC 9278 ± 853 1114 ± 23 1727 ± 52 637 ± 38 8868 ± 588

PF111-clear

216 EC 5540 ± 214 985 ± 39 2437 ± 72 901 ± 29 14843 ± 611

Table 5. Summary of spray-formed EMAA mechanical properties.
USACERL TR 98/14 49

Figure 20. Stress-strain response of PF111 when pigmented, clear, and quenched.
50 USACERL TR 98/14

Figure 21. Selected mechanical properties for PF111: a) modulus, b) tensile, c) elongation to break, and

d) toughness.
USACERL TR 98/14 51

Figure 22. Stress-strain response for PF113 when pigmented, clear, and quenched.
52 USACERL TR 98/14

Figure 23. Selected mechanical properties for PF113: (a) modulus, (b) tensile, (c) elongation to break, and

(d) toughness.
USACERL TR 98/14 53

Recycled Blend

(PCCP:PF111)

Secant

Modulus at

5 percent strain

(psi)

Yield Strength

(psi)

Tensile

Strength

(psi)

Elongation to

Break ( percent)

Toughness

(lb/inch)

10 percent PCCP 12283 ± 620 1279 ± 81 1296 ± 81 114 ± 18 1254 ± 277

20 percent PCCP 16259 ± 1355 1580 ± 83 1592 ± 98 62 ± 12 8777 ± 159

30 percent PCCP 19283 ± 998 1758 ± 76 1760 ± 80 31 ± 3 446 ± 60

40 percent PCCP 200992 ± 1091 1780 ± 49 1783 ± 50 25 ± 3 343 ± 54

50 percent PCCP 23986 ± 1225 1944 ± 74 1946 ± 46 19 ± 3 271 ± 35

60 percent PCCP 28644 ± 881 2176 ± 78 2179 ± 75 14 ± 1 219 ± 17

70 percent PCCP 32724 ± 1973 2269 ± 42 2270 ± 43 11 ± 0.5 160 ± 13

80 percent PCCP 34544 ± 868 2230 ± 199 2230 ± 200 8 ± 1 144 ± 28

90 percent PCCP 37926 ± 1593 2273 ± 55 2274 ± 56 7 ± 1 103 ± 14

100 percent PCCP 42723 ± 2857 2126 ± 197 2126 ± 197 5 ± 0.5 62 ± 11

Table 6. Mechanical properties of melt-compounded PCCP with EMAA copolymer

Figure 24. Stress-strain response for PCCP blends; volume percent PCCP noted on plot.
54 USACERL TR 98/14

Figure 25. Mechanical properties of PCCP blends: (a) yield strength, (b) tensile strength, (c) secant

modulus, and (d) elongation to break.
USACERL TR 98/14 55

Figure 25. Continued.
56 USACERL TR 98/14

Figure 26. Summary of peel strength obtained for varous steel surfaces.

Figure 27. Load-displacement plot for PF111 on polished steel.
USACERL TR 98/14 57

20 μm

a)

b)

200 μm

bands

band

width

Figure 28. Polymer fracture surface from polished steel depicting bands at (a) low and (b) higher

magnification.
58 USACERL TR 98/14

Figure 29. Load-displacement plot for PF111 on polished steel.
USACERL TR 98/14 59

a)

b)

200 μm

200 μm

Figure 30. Peeling fracture surfaces from abrasive-blasted steel (a) PF111 and (b) PF113.
60 USACERL TR 98/14

polymer

20 μm

Figure 31. EMAA peeled from abrasive-blasted steel.
USACERL TR 98/14 61

Figure 32. (a) Load displacement plot for as-received steel surface and (b) SEM micrograph of polymer

coating surface.
62 USACERL TR 98/14

Figure 33. Proposed interfacial interactions between EMAA copolymer and steel oxide.
USACERL TR 98/14 63

4 Field Application

Based upon preliminary results of this work, the Triborough Bridge and Tunnel

Authority, New York, and SUNY Thermal Spray Laboratory personnel designated

two bridge test sites. The first demonstration took place during the third week of

August 1995 on Randall’s Island Pier 4. Approximately 1200 ft2

 was abrasiveblasted to a 2-3 mil SP-6 profile. Both virgin ethylene methacrylic acid copolymer

(PF111-grey) and a 3:1 blend of EMAA:PCCP were sprayed onto the designated

areas.

The second demonstration was completed with the Triborough Bridge and Tunnel

Authority during the third week of October 1995. Approximately 400 ft2

 was

abrasive-blasted to a 2-3 mil SP-6 profile. Both virgin EMAA copolymer (PF111-

grey) and a 3:1 blend of EMAA:PCCP were sprayed onto the designated areas. In

addition, adjoining bridge sections were painted with a two-part alkyd system for

comparison purposes.

In both test cases, the areas to be topcoated were preheated to 190 EF and a

minimum of 20 mils of thermoplastic were applied (Figure 34). Inspection after 10

months of exposure found no signs of deterioration at either test site.
64 USACERL TR 98/14

a)

b)

Figure 34. Field demonstration at the Triborough Bridge: (a) preheating the steel panel and (b) deposition

of recycled polymer blend.
USACERL TR 98/14 65

5 Cost Considerations

The addition of PCCP waste materials to virgin EMAA has the potential to reduce

the materials cost of flame-sprayed polymer coating systems. Savings will be

proportional to the volume of PCCP used. Using a 25 percent volume fraction of

PCCP, net savings after processing costs is estimated at 15–20 percent below the

cost for 100 percent virgin EMAA.

The total costs for field painting of steel structures are largely determined by the

complexity of site setup, the degree of surface preparation required, and any

requirements for highly skilled operators and specialized equipment. Materials

comprise only a small portion of the total cost for surface preparation and thermal

spray coating application. The percentage of total painting costs attributable to

coating materials ranges from about 20 percent for a simple steel structure to less

than 10 percent for structures that are intricate or difficult to access. Therefore, it

can be seen that a 20 percent reduction in materials costs does not represent a

significant savings in the total cost of a coating project, and this reduction is largely

offset by additional costs of surface preparation and specialized equipment.

A number of low-VOC conventional paint systems are available that will adequately

protect steel in atmospheric exposures. These can be applied with conventional

paint spray equipment by workers with a moderate skill level. Many such coatings

will tolerate an abrasive-blasted surface cleaned to a commercial or near-white

grade at much lower cost than the white metal surface preparation grade

recommended for flame-sprayed polymer coating systems.
66 USACERL TR 98/14

6 Conclusions, Recommendations, and

Commercialization

Conclusions

A design of experiments approach was used to investigate the effects of material

and application parameters on the melting and coalescence of a low molecular

weight EMAA copolymer and high molecular weight PCCP. The substrate preheat

temperature showed the largest effect on coating temperature. The powder

injection rate had the largest effect on coating surface roughness and splat

elongation ratio. The results of this study indicate that thermal sprayed

EMAA/PCCP coatings should be applied only to surfaces that have been abrasiveblasted to the white metal grade.

Microstructures of the as-sprayed PCCP coatings (from dry-compounded powder)

showed evidence of cracking. Accelerated weathering tests induced further

cracking, photo-bleaching, and greater surface roughness.

The dry-compounded PCCP powder was blended with virgin EMAA in volumetric

proportions of 25, 50, and 75 percent. Test results showed that the blending of

EMAA with PCCP decreases the weathering-induced crack density and the average

crack length as compared with unblended PCCP. No cracking was evident in the

100 percent EMAA coatings.

Blending the EMAA copolymer with PCCP also increased the adhesion to

abrasive-blasted steel from 250 psi for 100 percent PCCP to 1100 psi for 25 percent

PCCP. The 1250 hours of accelerated QUV weathering did not appreciably affect

coating adhesion. Melt-compounded PCCP has greater strength and a higher

modulus in comparison to virgin EMAA. However, coating toughness can be

improved by incorporating EMAA into the feedstock.

In summary, it is concluded that the incorporation of PCCP into virgin EMAA

material is associated with some loss of performance. Cost savings resulting from

the blending 25 percent PCCP with 75 percent virgin EMAA are insignificant
USACERL TR 98/14 67

compared to the total cost of a recoating job. Therefore, it i s expected that the use

of PCCP/EMAA will have only a very limited commercial usefulness.

Recommendations

The use of PCCP/EMAA is not recommended as a protective coating for steel

structures.

Commercialization and Technology Transfer

A draft Civil Works Guide Specification (CWGS) was to have been prepared for this

technology upon conculsion of the work. However, the unremarkable performance

of PCCP/EMAA blends in this investigation indicates that the preparation of such

a specification would be highly premature.

The State University of New York continues to monitor coating performance at the

two field demostration sites and conduct related research. If the SUNY researchers

are able to overcome the technical problems shown in this study, they will initiate

work with the Steel Structures Painting Council and other standards organizations

to develop materials standards and application guidance.

The State University of New York maintains responsibility for the

commercialization of the polymer coatings. SUNY has undertaken an extensive

technology transfer program to promote commercialization of this technology. To

date this program has included the activities noted below.

Patent Disclosure

J. Brogan, C.C. Berndt, R. Lampo, K.A. Gross, S. Sampath, and H. Herman,

Thermal Spraying of Recycled and Post-Consumer Commingled Polymer

Material, S-7088, disclosed 17 November 1994.

Invited Presentations

1. A Comprehensive Approach Towards the Maintenance of Infrastructural Steel,

Northwestern University Infrastructure Technology Institute, Absecon, NJ,

1–3 November 1995.

2. The Role of Academia: Strategic Partnerships for Advancing Thermal Spray

Technology, National Thermal Spray Conference (NTSC '95), Thermal Spray

Society, 10 September 1995.
68 USACERL TR 98/14

3. Presentation to the Welding Research Council Subcommittee on Hardfacing

and Wear, 21 April 1995.

Conferences

1. 4th World Congress on Coating Systems for Bridges and Steel Structures, 1-3

February, 1995, St. Louis, MO (two presentations)

2. National Association of Corrosion Engineers, 27-31 March, 1995, Orlando, FL

(one paper presented)

3. American Welding Society, 3-7 April, 1995, Cleveland, OH (one paper

presented)

4. National Association of Corrosion Engineers, 23-28 March, 1996, Denver, CO

5. 1996 VIII International Congress on Experimental and Applied Mechanics,

June 10-13, Nashville, TN (one paper presented)

6. ASM International, 5th National Thermal Spray Conference, May 1994,

Boston, MA (eight papers)

7. ASM International, 6th National Thermal Spray Conference, 11-15 September

1995, Houston, TX (six papers)

8. Civil Engineering Research Foundation (CERF) Meeting on Infrastructure,

1-3 April 1995, Washington, DC.

Selected Publications

1. S.H. Leigh and C.C. Berndt, “A Test for Coating Adhesion on Flat

Substrates— A technical note,” JTST 3[2] (1994), pp 184-190.

2. T. Sugama, R. Kawase, C.C. Berndt, and H. Herman, “An Evaluation of

Methacrylic Acid-modified Polyethylene Coatings Applied by Flame-spray

Technology,” Progress in Organic Coatings, 25[2] (1995), pp 205-216.

3. J.A. Brogan and C.C. Berndt, “The Coalescence of Combustion Sprayed

Ethylene-Methacrylic Acid Copolymer,” accepted for J.Mater. Sci., April 1996.

4. J.A. Brogan, C.C. Berndt, W.C. Smith, R.V. Gansert, S. Raghu, S. Sampath,

and H. Herman, “Real-time Imaging of the Plasma Spray Process—work in

Progress,” JTST, 4[4] (1995), pp 374-376.

5. J.A. Brogan, K.A. Gross, Z. Chen, H. Herman, and C.C. Berndt, “Investigation

of Combustion Sprayed Hydroxyapatite/polymer Composite Coatings,”

pp 159-164 in 1994 Thermal Spray Industrial Applications, ed., C.C. Berndt

and S. Sampath (ASM International, Materials Park, OH, 1994).

6. J.A. Brogan, R. Lampo, and C.C. Berndt, “Thermal Spraying of Polymers,”

Proc. 4th World Congress on Coating Systems for Bridges and Steel Structures,

1-3 February, 1995, St. Louis, MO, pp 200-212.
USACERL TR 98/14 69

7. J.A. Brogan, J. Margolies, S. Sampath, H. Herman, C.C. Berndt, and S.

Drozdz, “Adhesion of Combustion-Sprayed Polymer Coatings,” pp 521-526 in

1995 Advances in Thermal Spray Science and Technology, ed. C.C.Berndt and

S.Sampath (ASM International, Materials Park, OH, 1995).

8. C.C. Berndt, M.L. Allan, H. Herman, J. Brogan, R. Benary, R. Zatorski, and

M. Leote, “Thermal Spray Technologies for Infrastructure Repair and

Maintenance,” extended abstract for Proc. of the VIII International Congress

on Experimental Mechanics, Pub. SEM, CT-USA, 1996.

9. World Wide Web Home Page. An informational home page is located at

http://dol1.eng.sunysb.edu/berndt.html.
70 USACERL TR 98/14

References

1. T. Katauskas, “DOT Coats Rusting Bridges with Layers of Problems,”

Research and Development, May, (1990) pp 42-48.

2. T. Race, V. Hock and A. Beitelman, “Performance of Selected Metallized

Coatings and Sealers on Lock and Dam Facilities,” J. Protective Coatings &

Linings, 6 [1], Jan. 1989, p 37-44.

3. J.A. Apostolos, D.M. Parks, and R.A. Carello, “Cathodic Protection using

Metallized Zinc,” Materials Performance, 26 [12], (1987) p 22-28.

4. R.H. Unger, “Thermal Spraying of Bridges,” Thermal Spray: Advances in

Coating Technology, Ed. by D. Houck, ASM, (1988) p 399-406.

5. G.K. Sweet, “Applying Thermoplastic/Thermoset Powder with a Modified

Plasma System,” Thermal Spray Coatings: Research Design and Applications,

Ed. by C.C. Berndt and T.F. Bernecki, ASM, (1993) p 381-384.

6. P. J. Loustaunau and D. Horton, “EMAA Thermoplastic Powder Coatings in

Shop and Field Applications,” Materials Performance, July (1994) p 32-35.

7. J. Maty, “Reg-Neg Plan Likened to L.A. Rules,” American Paint and Coatings

Journal, 78 [8], 1993.

8. A. Rogerson, “CALTRANS Tests Powder Coatings in the Field,” J. Protective

Coatings and Linings, 11 [8] (1994) p 25-32.

9. T. Race, “Evaluation of Flame-Sprayed Polymer Coatings for Civil Works

Navigation Structures,” USACERL Technical Report FM-94/07, April 1994.

10. T.W. Glass and J.A. dePay, “Protective Thermoplastic Powder Coating

Specifically Designed Adhesive Polymers,” Thermal Spray Coatings:

Properties, Processes and Applications, Edited by T.F. Bernecki, ASM, (1991)

p 345-351.

11. J.A. Brogan, K.A. Gross, Z.Chen, C.C. Berndt, H. Herman, “Investigation of

Combustion Sprayed Hydroxyapatite/Polymer Composite Coatings,” Thermal

Spray Industrial Applications, Edited by C.C. Berndt and S. Sampath, ASM,

(1994) p 159-163.

12. Y. Bao and D.T. Gawne, “Plasma Spraying of Glass-Polyamide Composite

Coating,” Thermal Spray Coatings: Research Design and Applications, Edited

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13. J.A. Brogan and S. Reddy - 1994 Dupont Plunket Student Award.

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72 USACERL TR 98/14

Appendix: Equipment and Instrumentation

Vendors

Plastic Flamecoat Systems

1613-T Hwy 3

League City, TX 77573

(713)332-8180

Applied Test Systems, Inc.

PO Box 1529

Butler, PA 16003

(412)283-1212

Tinius Olsen Testing Machine Co.

Easton Road

Willow Grove, PA 19090

(215)675-7100

Stat-Ease, Inc.

2021-T E. Hennepin Ave, Suite 191

Minneapolis, MN 55413

(612)378-2152

Control Vision, Inc.

PO Box 51505

Idaho Falls, ID 83405

(208)523-5506

Buehler Ltd.

41 Waukegan Road

Lake Bluff, IL 60044

(800)283-4537

Unitrol, Inc.

170-T Wilbur Place

Bohemia, NY 11716

(516)589-6666

Obex, Inc.

PO Box 1253

Stamford, CT 06904

(203)975-9094

Perkin-Elmer Corp.

Metco Division

PO Box 1006

Westbury, NY 11590

(516)334-1300
USACERL TR 98/14 73

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