Dynamic densiﬁcation of metal matrix-coated ﬁbre

Acta Materialia 53 (2005) 617–628

www.actamat-journals.com

Dynamic densification of metal matrix-coated fibrecomposites: modelling and processing

H.X. Peng a,1, F.P.E. Dunne b,*, P.S. Grant a, B. Cantor a,2

a Oxford Centre for Advanced Materials and Composites (OCAMAC), Department of Materials, Oxford University, Parks Road, OX1 3PH, UKb Department of Engineering Science, Oxford University, Parks Road, Oxford OX1 3PJ, UK

Received 4 March 2004; received in revised form 17 August 2004; accepted 11 October 2004

Available online 11 November 2004

Abstract

The consolidation processing of Ti–6Al–4V matrix-coated fibre (MCF) composite under vacuum hot pressing (VHP) has been

investigated. A new test methodology has been developed for the determination of in situ matrix coating creep properties. In using

the methodology, only a single, simple test is required, together with finite element modelling of the single fibre compression test.

The creep coefficient and stress index have been determined for electron beam evaporated physical vapour deposited Ti–6Al–4V at

900 �C to be 1.23 · 10�5 and 1.3, respectively. Consolidation experiments have been carried out on multi-ply MCF arrays under

vacuum hot pressing. Finite element models have been developed for the dynamic consolidation of both square and hexagonal fibre

packings. The creep constants for the Ti–6Al–4V, determined using the single fibre test, were assigned to the coating in the finite

element models. Excellent agreement between predicted and experimental results was achieved, providing verification of the single

fibre test methodology for the determination of creep constants.

� 2004 Acta Materialia Inc. Published by Elsevier Ltd. All rights reserved.

Keywords: Single fibre tests; Matrix-coated fibres; Vacuum hot pressing; Dynamic densification; Titanium matrix composites

1. Introduction

Titanium matrix composites (TMCs) in which the

reinforcement is continuous silicon carbide fibre are

being developed for application in aircraft engine com-

ponents such as compressor bladed rings (blings). TMCs

offer superior properties compared to conventionalmaterials and may allow significant reduction in compo-

nent weight. Among the techniques developed for man-

ufacturing TMCs [1,2], the matrix-coated fibre (MCF)

1359-6454/$30.00 � 2004 Acta Materialia Inc. Published by Elsevier Ltd. A

doi:10.1016/j.actamat.2004.10.015

* Corresponding author. Tel.: +44 1865 273 140; fax: +44 1865 273

905.

E-mail address: [email protected] (F.P.E. Dunne).1 Present address: Department of Aerospace Engineering, Univer-

sity of Bristol, University Walk, Bristol BS8 1TR, UK.2 Present address: Vice-Chancellor�s Office, University of York,

Heslington, York YO10 5DD, UK.

method offers a number of advantages over others, such

as uniform fibre distribution, reduced fibre damage and

better fibre fraction control during consolidation

processing [3–5]. In the MCF method, the fibres are

packed into arrays and consolidated either by hot iso-

static pressing (HIP) or by vacuum hot pressing (VHP)

to provide the finished TMC. Long fibre-reinforced me-tal matrix composites are unlikely to be used in their

own right for component manufacture. Instead they

are likely to be used as local reinforcement in a larger

metal structure. In this context, matrix-coated fibres

can be used to produce locally reinforced TMC compo-

nents and the VHP technique is likely to be more prac-

tical than HIP because it allows the consolidation of the

MCFs directly into a die of desired component material.Thus, the consolidation of MCFs via the VHP technique

is of much interest [6].

ll rights reserved.

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618 H.X. Peng et al. / Acta Materialia 53 (2005) 617–628

The consolidation process is important in determin-

ing the properties of the final composite. Pressure and

temperature cycles have to be optimized to ensure the

quality of the final product [1]. In order to optimize

these parameters, theoretical simulations that can pro-

vide an insight into the fabrication process [7–15] andsome experimental studies [6,10] have been carried out.

However, there are two problems with existing work.

First, the material properties used in the models are of-

ten taken from bulk matrix materials such as Ti–6Al–

4V. The properties of bulk material are not the same

as those of the Ti–6Al–4V coating produced during

MCF manufacture by physical vapour deposition

(PVD). Although PVD bulk materials have been pro-duced and investigated to provide improved property

data [16], the properties of the unique microstructure

of the Ti matrix formed during PVD onto fibres are

not well established [17,18]. Second, while many of the

theoretical simulations were carried out by assuming a

square fibre array and only a few studies have consid-

ered the hexagonal array [9,10], the experimental data

were mostly acquired from randomly packed fibres.Akisanya et al. [7] employed steel cylinders with diame-

ters of 12 and 8.2 mm coated with plasticine with thick-

ness of 1.5 and 3.4 mm as an alternative in their

experimental studies to simulate a square array of

MCFs. This approach was adopted because of the tech-

nical complexity and difficulty in achieving a regular fi-

bre arrangement in MCFs. Due to the absence of

regular fibre packing, both matrix deformation and fibrerearrangement occur during the consolidation process.

Consequently, the fibre arrangement has not been well

controlled to provide valid experimental results for ver-

ifying theoretical predictions.

Existing modelling work has provided valuable in-

sight into the consolidation process, and provides, for

example, stress–strain data during densification, the size

and shape of the yield surface [1,8], and details of thematerial flow [10]. Carmai and Dunne [19,20] have

developed physically based constitutive equations for

the consolidation of MCFs which require just two mate-

rial properties; namely, the creep coefficient and expo-

nent for the titanium matrix coating. The model,

implemented within the FE method, is able to predict

the consolidation behaviour given the process parame-

ters of pressure and temperature, and details of the par-ticular composite selected (fibre size and volume

fraction) [19]. A further simplified model has also been

developed for simulating practical consolidation proc-

esses [21], which also relies on knowledge of the matrix

creep coefficient and exponent.

In the present work, a novel test method is developed

to determine the MCF coating creep properties. This is

achieved by carrying out compressive tests on a singlematrix-coated fibre, under transverse loading. The

extraction of the matrix creep properties relies on finite

element modelling of the single-fibre compression test.

The single-fibre test methodology presented provides a

comparatively simple way to obtain matrix coating

properties by testing a single-coated fibre. Constrained

uniaxial consolidation tests are also carried out in which

square and hexagonal fibre packings are carefully main-tained, using test methods already developed [22]. These

tests are also modelled using FE techniques, using the

creep properties obtained from the single-fibre tests.

The FE predictions are compared with the results of

the experiments, so that the effectiveness of the single fi-

bre test in determining correctly the creep properties can

be established.

The dynamic densification behaviour of MCFs dur-ing uniaxial vacuum hot pressing is then investigated

in which dies made of Ti–6Al–4V alloy have been used

to generate conditions closer to those experienced in

practical consolidation processing. We investigate both

fibre rearrangement and microstructural evolution in

processing, and present results of scientific and practical

interest.

2. Single fibre testing

2.1. Experimental arrangement

Ti–6Al–4V matrix-coated SiC monofilament fibres

with a diameter of �0.245 mm and SiC volume fraction

of �35% were supplied by Rolls-Royce plc, Derby, UK.The single fibre compression tests were conducted using

a vacuum hot press. The experimental arrangement is

shown schematically in Fig. 1. A number of fibres were

positioned parallel to each other between two precisely

machined ceramic discs with diameter of 50 mm and sur-

face roughness of �0.2 lm. The fibres were well spaced

relative to one another to ensure that they did not touch

during the compression testing. The length of each fibrewas no less than 100 times that of its diameter, so that

plane-strain conditions applied. Two identical steel

plates were heated by symmetric induction coils. To as-

sure the highest possible uniformity of compression of

the fibres, a ceramic ball was placed under the hydraulic

ram. The arrangement was centralized along the axis of

the ram. To measure deformation of a few tens of mi-

crons, a specifically designed and manufactured hightemperature extensometer together with its signal condi-

tioner (Model 3448 RCIH-010-015, Epsilon Technology

Co., Jackson, USA) and a data logging unit with 16 bit

resolution were employed. This enabled a deformation

of 0.25 lm to be repeatedly resolved within a 10 mm

gauge length. The extensometer was enclosed in a

water-cooled box with two ceramic rods attached to

the upper and lower ceramic discs where the tempera-ture was monitored using K-type thermocouples. After

pulling the vacuum, the temperature was ramped at a

Fig. 1. Experimental set-up for the single-fibre compression testing.

Upper Movable boundary

Fibre

Matrix

Ux=Uy=0

Lower unit-cell boundary

Lef

tu

nit

-cel

lbo

un

dar

y

Fig. 2. Finite element model developed for simulating single-fibre

compression.

H.X. Peng et al. / Acta Materialia 53 (2005) 617–628 619

heating rate of �10 �C /min to 900 �C and held for

15 min to allow thermal equilibrium to be attained

and the readings from the extensometer to become sta-ble. A constant load was then applied and the nominal

pressure was calculated according to the length and

the number of fibres. The data logger recorded the ver-

tical displacement versus time. After testing, the de-

formed fibres were cross-sectioned, mounted and

polished. Both polished and etched samples were exam-

ined using optical microscopy with DIC/polarized light.

2.2. Modelling the single fibre compression

In the finite element model developed, it was assumed

that the SiC fibre was stiff compared to the titanium al-

loy matrix, especially at elevated temperatures. The elas-

tic moduli of SiC and Ti–6Al–4V at 900 �C were

assumed to be 354 and 28.4 GPa, respectively [10].

The SiC fibre was therefore assumed to be rigid. Thematrix coating was considered to be incompressible

and to deform by power-law creep. Uniaxial power-

law creep is given by

_e ¼ Arn; ð1Þwhere _e is creep rate, r the stress, A the creep coefficient

and n the stress exponent. This can be generalized intomultiaxial form as

_eij ¼3

2Arn�1

e rij; ð2Þ

re denotes the effective stress defined by

r2e ¼

3

2r0ijr

0ij; ð3Þ

where r0ij ¼ rij � dijrm is the deviatoric stress, rm the

mean stress and dij is the Kronecker-Delta symbol.

The constitutive power-law creep equation (2) was

implemented into ABAQUS finite element software

using a UMAT subroutine.

The finite element model developed for the single fi-bre compression using ABAQUS is shown schematically

in Fig. 2. Due to symmetry, one quarter of the fibre was

modelled. Since the fibre was assumed to be rigid, the fi-

nite element mesh was only developed for the matrix

coating which consisted of 2D, four-noded plane strain

elements. All nodes lying on the fibre-matrix boundary

were fixed in both the x and y directions and those on

the lower unit-cell boundary were allowed to move inthe x-direction only while all nodes lying on the left

unit-cell boundary were allowed to move in the y-direc-

tion only. An upper movable boundary was used for the

application of the distributed load simulating the pres-

sure state during the compression process and was al-

lowed to move in the direction of the y-axis. Sticking

friction was assumed between the matrix and the mova-


ble boundary. Examination of tested specimens and

comparison with finite element modelling results showed

this to be a good assumption at 900 �C. The pressure

applied on the movable boundary was in the range of

20–40 MPa. The displacement of the upper movable

boundary simulated the extensometer-measured dis-placement in the experiments. By choosing values for

the creep constants A and n, finite element simulations

of the test could be carried out. By carrying out para-

metric finite element studies over ranges of values for

A and n, and by comparing the resulting predicted dis-

placements with experimental measurements, the opti-

mum values of A and n to give the best fit to the

experimental data was obtained. For a particular nomi-nal pressure of 30 MPa, the calculated and experimen-

tally determined displacements are shown in Fig. 3(a)

for the optimum creep constants. In addition to the

extensometer-determined displacements, the displace-

ment was also measured by using an image analyser

Fig. 3. (a) Best fit finite element displacement for the single MCF

compression test at 900 �C and 30 MPa and (b) micrograph of the

deformed MCF with superimposed finite element mesh.

and micrographs obtained under dark mode in which

the coating edges were highlighted. Micrographs were

obtained by carrying out repeat tests to different levels

of displacement. These results are also shown in Fig.

3(a) by the symbols. The predicted and experimentally

determined single matrix-coated fibre deformations atthe end of the test are shown in Fig. 3(b), and good

agreement was obtained.

It is argued that the values of A and n so obtained are

the true creep constants for the coating and that they are

unique. This is because the creep deformation occuring

in the coating is highly inhomogeneous, with heteroge-

neity in strain, stress, and importantly, strain rate. In

some ways, this can be considered similar to carryingout many homogeneous, uniaxial tests on the coating

over a wide range of stress level. This is the conventional

way to identify creep constants. The optimum values of

the creep constants A and n for 900 �C were found to be

1.23 · 10�5 and 1.3, respectively. In the subsequent sec-

tions, we examine in detail whether these are correct. If

it can be demonstrated that they are, then it is argued

that we have developed a simple testing methodologyfor the determination of the creep properties of the fibre

matrix coating.

3. Investigating the validity of the test methodology

In order to test the correctness, or otherwise, of the

creep properties obtained, experimental tests were car-ried out on both square- and hexagonal-coated fibre

packing arrays under conditions of constrained uniaxial

consolidation. Both cases were also modelled, and the

matrix coating assigned the creep constants determined

above. The correct prediction of the consolidation

behaviour in the two cases provided evidence that the

creep properties are correct.

The fibre square and hexagonal arrays were ob-tained by firstly winding fibre onto an aluminium

drum to form a ply with 24 fibres. A small amount

of acrylic binder was used to hold the fibres together.

The binder was subsequently burned off at �320 �C.The fibre-ply were cut into the desired length (no less

than 100 times that of the fibre diameter) and posi-

tioned into a hot press die made of nickel-alloy as

shown in Fig. 4. The advantage of this die assemblywas that it allowed a fine control of the fibre arrange-

ment by varying the width of ram �D� according to the

diameter and number of MCFs in the fibre-ply, since

the fibre sideways motion was well constrained at both

sides. More details can be found elsewhere [22]. The

extensometer was attached on the upper and lower

ram �D�. Due to the limitation on the gauge length

and full-scale travel of the extensometer, the maximumnumber of fibre layers used in these experiments was

Fig. 5. Finite element model developed for square fibre array.

Fig. 6. (a) Micrograph showing a square fibre array early in the

consolidation process; (b) the deformed finite element mesh superim-

posed on the micrograph showing excellent agreement; and (c) the

resulting fully dense rectangular fibre array obtained under contrained

uniaxial compression.

Fig. 4. Die assembly for MCF consolidation experiments.


10. To make a square array, the number of fibres in

each layer was the same; 24 in this case. For a hexag-

onal array, the number of fibres was alternated layer

by layer with a difference of one fibre in each layer,

that is, with fibre numbers of 24/23/24/23� � �24/23.The fibre arrangement had a significant effect on the

measured dynamic densification behaviour, in particu-

lar, the displacement versus time curve. In addition,the two arrangements led to quite different initial rela-

tive densities; hexagonal packing provided the higher

initial relative density.

The vacuum chamber was purged with Ar followed

by evacuation. The die assembly was then heated with

a ramping rate of �10 �C/min. After the temperature

reached 900 �C and the extensometer reading became

stable, a constant load was applied for a specified time.At the end of the test, the induction power was cut off

and the die assembly was subjected to natural cooling

in vacuum. Cooling to 300 �C took �2 h. A data logger

recorded displacement, load and temperature. The

cross-sectioned sample was polished and examined using

optical microscopy with DIC/polarized light to reveal

the microstructure of the composite, with particular

attention to fibre arrangement.

3.1. Consolidation of MCFs with square array

configuration

The finite element model developed to simulate uni-

axial constrained consolidation of multiple fibres ar-

ranged with square packing is shown schematically in

0

100

200

300

400

0 500 1000 1500 2000

Time / s

Dis

pla

cem

ent

/m

0

400

600

800

1000

Lo

ad /

kN, T

emp

erat

ure

/ oC

FEM-disp.

Temp.

E XP-disp.

Load 5

Fig. 7. Predicted and experimental displacement versus time curves

obtained for the consolidation of a square fibre array under

constrained uniaxial compression at 900 �C and 30 MPa.


Fig. 5. For reasons of symmetry, only one quarter of the

coated fibre was modelled. The right non-movable

boundary was also a line of symmetry to produce a

repeating cell in which no fibre displacement in the hor-

izontal direction was allowed, simulating the con-

strained uniaxial compression. Pressure was applied to

the upper movable boundary, as in the experiments.

Sticking friction was assumed between the matrix andthe upper movable boundary because of the rapid effects

of diffusion bonding when the titanium alloy coating of

one fibre comes into contact with another. The friction

condition between the matrix and the right non-movable

boundary had very little effect on the predicted results

and hence no friction was assumed. The creep constants

assigned to the fibre coating were those determined

above from the single fibre test.Fig. 6(a) shows a micrograph of the square packed fi-

bre array early in the consolidation process. Note the

good control of fibre packing. Fig. 6(b) shows a closing

void in more detail together with the excellent agreement

with the superimposed finite element model prediction

Fig. 8. Finite element model develop

at the corresponding time in the consolidation process.

On completion of consolidation, the voids were com-

pletely removed as seen in Fig. 6(c). Due to the con-

strained uniaxial compression, the centre to centre

distance between fibres in the vertical direction was less

than that in the horizontal direction, with the latterremaining almost unchanged. This results in the near-

perfect rectangular fibre array at the end of the process,

shown in Fig. 6(c). The experimental and predicted

overall vertical displacements versus time are shown in

Fig. 7, and again there was excellent agreement, which

was reproduced for similar experiments carried out at

20 and 40 MPa. However, the sample consolidated at

20 MPa for 50 min was not fully dense at the end ofthe test indicating that this pressure was too low and,

consistent with Schuler�s work [10], suggested a thresh-

old value of �30 MPa for densification in reasonable

times.

3.2. Consolidation of MCFs with hexagonal array

configuration

The finite element model developed for hexagonal fi-

bre packing is shown schematically in Fig. 8. The repre-

sentative unit cell did not apply for the fibres near to the

vacuum hot press die. This would be likely to lead to dif-

ferences in predicted and experimental results if a small

number of fibres were used in the experiments. However,

because each layer in our experiments contained 23 or

24 fibres, this was unlikely to be relevant. Sticking fric-tion was assumed between the contacting fibre coatings.

The upper movable boundary was a rigid surface that

was allowed to move only in the vertical direction. A

force was applied to the rigid surface to generate the re-

quired applied pressure. The creep constants assigned to

the fibre coating were, again, those determined above

ed for a hexagonal fibre array.

Fig. 9. Micrograph showing a consolidated hexagonal fibre array with

finite element model superimposed, showing good agreement.

0

30

60

90

120

150

180

0 200 400 600 800 1000 1200 1400Time / s

Dis

pla

cem

ent

/ µm

0

200

40

600

800

1000

Lo

ad /

kN, T

emp

erat

ure

/ oC

Tem

Exp.-

FEM-

Load 5

Fig. 10. Predicted and experimental displacement versus time curves

for a hexagonal fibre array under constrained uniaxial compression at

900 �C and 30 MPa.


from the single fibre test. The results of a constrained

uniaxial compression test carried out at 900 �C and at

30 MPa on a hexagonally packed fibre array are shown

in Fig. 9, at the end of the experiment, during which thehexagonal packing was carefully maintained. Superim-

posed on the micrograph is the finite element model pre-

diction for the corresponding time, showing good

agreement. The predicted and experimental overall ver-

tical displacement versus time curves are shown in

Fig. 10, and again, good agreement between prediction

and measurement is achieved. Comparing Figs. 10 and

7 shows that the time required to achieve fully densematerial was shorter for the hexagonal packing (�1200

s compared to �1500 s for the square array), and of

course that the displacement required to achieve this

was smaller (approximately half of that required for

the square array).

4. Effect of fibre rearrangement and microstructural

evolution in consolidation processing

4.1. Fibre rearrangement

It has been demonstrated that the fibre arrangementcan be maintained in either the square or hexagonal

packing formats which allows the dynamic densification

behaviour of MCF arrays to be extracted. This is mainly

attributed to the well-constrained uniaxial compression,

i.e. sideways motion of the fibres was prohibited by fine

control of the die assembly. The negligible deformation

of the nickel alloy dies in the temperature and pressure

ranges used in the present study also helped to retainthe fibre array.

In a practical consolidation process, however, fibre

rearrangement may inevitably occur due partially to

the die deformation at high temperature and hence less

constraint imposed on the fibre sideways motion [22].

In addition, the MCFs diffusion bond themselves to the

die in which they are being consolidated. To demonstrate

this, the consolidation of the MCFs into a titanium alloy(Ti–6Al–4V) die has been carried out. The die assembly

was heated up with a ramping rate of about 10 �C/min

to 900 �C. After compression of the MCF arrays for a gi-

ven time, the die assembly was allowed to cool naturally

for 2 h to 300 �C in vacuum. In order to obtain insight

into the deformation process, several experiments were

interrupted at different times and therefore extents of

consolidation. These samples were sectioned, mountedand polished. Precautions were taken to ensure the sur-

faces perpendicular to the fibre axes were polished to

the highest standards. Both polished and etched samples

were examined using optical microscopy with DIC/po-

larized light. Because the tight fit of the fibre-die assem-

bly at room temperature, the initial fibre arrangement in

the die was likely to be constrained to square rather than

hexagonal array packing, and this was confirmed byexamining the MCF array at the initial stage of the con-

solidation process as shown in Fig. 11(a). As consolida-

tion proceeded, it was reasonable to assume that

rearrangement and deformation co-existed as the fibre

array tended to the closest packed state – a hexagonal ar-

ray. The cross-section of a near-fully consolidated TMC

is shown in Fig. 11(b) comprising a fibre array with much

strengthened hexagonal character. Complete diffusionbonding between the matrix coating and the titanium al-

loy die is shown in Fig. 11(c).

The dynamic densification behaviour for a practical

fibre arrangement of mixed square and hexagonal

characters is shown in Fig. 12 together with those

for carefully controlled square and hexagonal fibre ar-

rays under identical processing conditions of 900 �Cand 30 MPa. The ram displacement was normalizedby the measured total displacement for each particular

fibre array. To densify a square array required a larger

Fig. 11. The cross-section of the fibre array: (a) at the initial stage of the consolidation process showing a near square fibre array; (b) the near-fully

consolidated composite for a near hexagonal fibre array; and (c) sound bonding with upper die at the initial stage.

0

0.2

0.4

0.6

0.8

1

0 400 800 1200 1600

Time / s

No

mo

rlis

ed D

isp

lace

men

t

Fig. 12. The normalized displacement versus time curves for a

practical process (heavy line), square fibre array (square symbol) and

hexagonal fibre array (diamond symbol) under identical processing

conditions.


displacement (typically, twice) than for a hexagonal ar-

ray because of the different initial packing density.

From Fig. 12, for the practical case where fibre rear-

rangement occurs, dynamic densification behaviour

falls between those for hexagonal and square array

packings. The vertical displacement (Dh) caused by

the fibre rearrangement during the transition from asquare to hexagonal array in uniaxial constrained

compression is [22]:

Dh ¼ ðN � 1Þð1� coshÞd; ð4Þ

where N is the number of fibre-layers, d is the fibre diam-

eter and h is the angle representing a particular stage of

the rearrangement, where 0 6 h 6 30�. The fibres form asquare array with h = 0 and a hexagonal array at

h = 30�. Fibre rearrangement is expected to occur

throughout the consolidation process and the value of

h at a particular stage is hard to quantify.


4.2. Microstructural variation during consolidation of

regular fibre arrays

Fig. 13 shows the microstructure of a fully consoli-

dated composite with �square� fibre array and with com-

plete diffusion bonding. Because of the largerdeformation along the vertical direction, a near equi-

axed matrix structure was formed in the region between

each pair of vertical fibres in Fig. 13(a), while a colum-

nar structure was retained between each pair of horizon-

tal fibres in Fig. 13(b). After experiencing the thermal

history described above, the matrix structure consisted

of a coarse a-plate and b-phase columnar structure. As

discussed later, the break up of the lamellae in the regionbetween each pair of vertical fibres was likely to be due

to the larger deformation compared with that between

each pair of horizontal fibres. The microstructure of a

hexagonal array deformed at 900 �C and 30 MPa for

only �5 min is given in Fig. 14(a). The regular fibre ar-

ray was near perfectly retained during consolidation.

Microstructural coarsening probably occurred during

the slow cooling process from the consolidation temper-

Fig. 13. The matrix microstructure with a square fibre array showing

(a) near-equiaxed structure between each pair of vertical fibres and (b)

columnar structure between each pair of horizontal fibres (the pressing

direction is vertical).

Fig. 14. Micrographs of the hexagonal fibre array: (a) at the start of

consolidation showing coarsening of the lamellar structure; (b) just-

consolidated composite showing incomplete diffusion bonding; and (c)

further-consolidated material showing improved diffusion bonding.

ature. Fig. 14(a) shows that the contact area betweeneach pair of �vertical� fibres was larger than that between

the horizontal fibres. The microstructure of a �just� fullyconsolidated sample (as indicated by the horizontal re-

gion of the displacement versus time curve on the data

logger screen) is shown in Fig. 14(b). All voids were

eliminated, and the relative density of the composite

was unity. Again, the fibre array was well maintained.


A particular feature of the microstructure shown in Fig.

14(b) was that the boundaries between adjacent fibre

coatings were still visible. This indicates that diffusion

bonding was not yet complete. Additional time at pres-

sure and temperature was necessary to ensure full bond-

ing. The microstructure of a sample compressed for anextra 8 min is shown in Fig. 14(c) where better bonding

was achieved. These findings may be of significant

importance for developing the manufacturing process

for a real component. Because of the reduced overall

matrix deformation required to consolidate a hexagonal

array, the break up of lamellae was less pronounced

than that observed in a square array as shown in Fig.

13(a). Nonetheless, some fine equiaxed a-phase was pre-sent at the triple points where the material free surfaces

impinged during the final stage of consolidation. Slight

a-phase refinement was also noted along the centre line

where the matrix free surfaces impinged.

Complete diffusion bonding in the square array oc-

curred at a comparatively early stage, in Fig. 13, because

of the high local contact stress between each pair of ver-

tical fibres. For a hexagonal fibre geometry, the localstress at the contacting area is expected to be lower than

that for square fibre arrays under equivalent external

load because of the larger number of vertical load-carry-

ing contacts per MCF.

4.3. Mechanisms responsible for the changes of the matrix

microstructure

The microstructural variation during the heat treat-

ment and hot working of Ti–6Al–4V with a lamellar

transformed microstructure has been a subject of a num-

ber of investigations [23–29]. The b transus temperature

above which the two phase a + b mixture transforms to

b is generally assumed to lie within the range of 995 �C[23] to 1015 �C [27], considerably above the consolida-

tion temperature of 900 �C used in the present study.Therefore, deformation of the Ti–6Al–4V matrix in the

MCFs takes place in the a + b two phase field. Research

on the thermomechanical processing of Ti–6Al–4V [28]

concluded that there was no dynamic (or metadynamic)

recrystallization within this temperature range. It has

been established reasonably well that the break up of

the a lamellae structure is caused by the globularization

or spheroidization, either static [29] or dynamic [30–32],of the colony structure. In view of the MCF consolida-

tion process, it is possible that both dynamic and static

globularization occurs during the pressing and cooling

processes. According to Weiss et al. [33], there are two

possible mechanisms for the breakup of the a lamellae

into low-aspect-ratio segments during hot deformation.

One is that both low and high angle boundaries across

the a-plates are formed, with misorientation angles var-ying from a few degrees up to about 30�. The b-phasecan then penetrate into the a-plates along these sub-

boundaries. If the a lamellae width is small, i.e. less than

two times the penetration distance of the b cusps, the

separation of a lamellae occurs readily. The penetration

rate depends on the diffusion of the alloying elements.

Another possible mechanism is that localized shear

and rotation of the a lamellae can occur during hotdeformation, and the misorientation across the shear

band can reach 20�. The b-phase may also penetrate

readily along the shear bands and partially or fully sep-

arate the a lamella into short segments. Other studies,

such as those by Ari-Gur and Semiatin [34] also con-

cluded that, at sub-transus temperatures, the globulari-

zation of a-plates did not occur by recrystallization

but by localized deformation of the plates and penetra-tion by the b-phase. Only moderate deformation was re-

quired to effect globularization whereas heavy

deformation normally yielded a large fraction of globu-

larization [33,34].

4.4. Dynamic densification behaviour of MCFs

Both experimental results and FE predictions indi-cate that a longer consolidation time is required for den-

sifying square fibre arrays than that for a hexagonal

array under the same conditions of temperature and

pressure. The typical dynamic densification curves (rela-

tive density against time) for consolidating both fibre ar-

rays under uniaxial constrained compression are very

similar to the curve shown in Fig. 12. It is suggested

that, for a practical process where fibre rearrangementmay occur, the densification curve falls between these

two curves, for example as indicated by the solid line,

in Fig. 12.

A high densification rate was observed at the begin-

ning of the vacuum hot pressing consolidation. Previous

investigations [10,35,36] showed an �S�-shaped densifica-

tion curve for the consolidation of a square fibre array by

hot isostatic pressing (HIP). These studies suggested thatdensification occurred in three stages as proposed by

Ward-Close and Loader [18]: an initial stage of low den-

sification rate as the consolidation conditions develop in

the HIP cycle; a second stage of rapid densification dri-

ven by high local contact stresses; and a final low densi-

fication rate as the final voids are consolidated. The low

densification rate at the initial stage was attributed to a

HIP encapsulating canister shielding effect, and simi-larly, a HIP canister used in consolidation also played

an important role in the densification of TMCs reported

by Vancheeswaran et al. [9]. In the present work, in the

absence of a canister, the first stage does not exist during

vacuum hot pressing and a rapid densification rate oc-

curred as a direct consequence of the high local contact

stresses at the initial consolidation stage. As consolida-

tion proceeded, for square fibre arrays, cusp-shapedvoids were formed and retained as shown in Fig. 11. This

is an important feature for consolidating MCFs because


micromechanical analyses of the creep collapse of cusp-

shaped pores indicated that their collapse rate could be

as much as an order of magnitude faster than that

of their spherical pore counterpart [37]. A recent study

of the effect of pore shape on densification concludes that

the densification rate of the material with circular poreswas much less than that of an identical material with

cusp-shaped pores [9]. Consequently, it is suggested that

the densification process can be made more time efficient

by retaining the initial cusp-shaped nature of pores.

5. Conclusions

The creep constants determined from the single fibre

test were used in the finite element predictions of the

densification of carefully controlled square and hexago-

nal fibre packing arrangements that showed quite differ-

ent densification time and displacement behaviour. In

general, excellent agreement between predicted and

experimentally determined results was achieved, and it

is concluded that the creep constants determined forthe fibre coating are therefore correct. We have therefore

been able to validate the single fibre test methodology

for the determination of fibre coating creep properties.

This offers advantages over alternative approaches in

that the test is simple to do, and that only one test is nec-

essary. In addition, the method enables the physical va-

pour deposited (PVD) coating itself to be tested, while

on the SiC fibre core, rather than an artificially gener-ated �bulk� Ti–6Al–4V material which typically has a dif-

ferent microstructure and physical behaviour to that of

Ti–6Al–4V on the SiC fibre.

The consolidation behaviour of matrix-coated fibre

composite under constrained uniaxial compression (vac-

uum hot pressing) has been investigated experimentally.

In contrast to most existing studies, square and hexago-

nal fibre arrangements have been achieved and main-tained during experiments by careful control of the die

assembly.

This control allowed the true dynamic densification

behaviour to be investigated because the effects of fibre

packing rearrangement were eliminated.

Finite element models based on a simple power law

constitutive equation describing the matrix deformation

have been developed and verified for simulating the con-solidation processing of matrix-coated fibres. Excellent

agreement between predictions and experiments was

achieved for both square and hexagonal fibre packings.

The Ti–6Al–4V coating deposited onto continuous

SiC monofilament fibre by PVD exhibited a basket-

weave microstructure containing very fine lamellar

a + b colonies. The a colonies coarsened rapidly during

the heating and cooling processes required for the con-solidation of MCF arrays into dense titanium matrix

composites. The break up of a lamellae and formation

of equiaxed �grains� occurred in regions of the MCF

coating where the largest deformation occurred because

of the static and dynamic globularization induced by the

localized deformation of a lamellae and the penetration

of b-phase.After the MCF reached full density, additional ther-

mal exposure under pressure was required to promote

complete diffusion bonding between adjacent MCFs,

particularly in the case of hexagonal fibre arrays.

Carefully controlled square-only and hexagonal-only

MCF arrays provided the lower and upper bounds,

respectively, for the dynamic densification behaviour

of MCFs under uniaxial, hot pressing. In the practical

case where MCF rearrangement occurred, the resultingdensification curve falls between these two bounds.

Acknowledgements

The authors are grateful to Rolls-Royce and the UK

Engineering and Physical Sciences Research Council for

support.

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Dynamic densiﬁcation of metal matrix-coated ﬁbre