THE DENSIFICATION OF METAL COATED FIBERS: … DENSIFICATION OF METAL COATED FIBERS: HOT ISOSTATIC PRESSING EXPERIMENTS ... Using a hot isostatic press containing an eddy current sensor

Actu maler. Vol. 45, No. 5. pp. IX51 ~1865. 1997 (‘I 1997 Acta Metallurgica Inc

Published by Elsevier S&nce Ltd Printed in Great Britain. All nghts reserved

1359.6454/97 $17.00 + 0.00

Pergamon PII: S1359-6454(96)00326-6

THE DENSIFICATION OF METAL COATED FIBERS: HOT ISOSTATIC PRESSING EXPERIMENTS

JOSEPH M. KUNZE and HAYDN N. G. WADLEYT Department of Materials Science and Engineering, School of Engineering and Applied Science,

University of Virginia, Charlottesville, VA 22903, U.S.A.

(Rewired 8 Ju1.v 1996; accepted 6 September 1996)

Abstract-An experimental study of the consolidation and associated grain growth kinetics of silicon carbide monofilaments sputter coated with nanocrystalline Ti-6A14V has been conducted. Bundles of coated fibers were consolidated using a hot isostatic press (HIP) equipped with an eddy current sensor for in situ measurement of the specimens’ relative density. The experiments were conducted at temperatures of 760”, 840’. and 900’C using a variety of pressure cycles that resulted in both fully and partially consolidated composites. The sensor data was used to determine the densification-time response of the specimens and to construct experimental HIP maps. Cross-sections of the partially consolidated specimens were used to investigate the evolution of coated fiber-fiber contacts and interfiber void shapes. The analysis indicated that the initial packing density was approximately random with an average of only about two contacts per coated fiber at a relative density of 70%. The number of contacts was a linear function of relative density reaching a maximum of almost six contacts per coated fiber at the theoretical density. The pores were found to be cusp-shaped throughout the consolidation process, indicative of a densification rate that greatly exceeded the (pore rounding) rate of sintering. H; 1997 Acta Metallurgica 111c.

1. INTRODUCTION

The high specific stiffness, strength, and creep resistance of fiber reinforced metal matrix composites (MMCs) that utilize either titanium or nickel alloy matrices and continuous silicon carbide monofilaments has attracted considerable aerospace interest [l]. In the past, many processing methods have been investigated for their production including a powder cloth approach (and its continuous tape casting analogue) [2], the foil-fiber-foil method [3, 41 and plasma spray deposition [S]. Recently, an alternative physical vapor deposition (PVD) process has emerged for composite preform fabrication [ 1, 6-101. In this approach, the matrix alloy is evaporated (either by sputtering or electron beam heating) and directly deposited onto the fibers [l, &lo]. This PVD synthesis route results in Sic monofilaments coated with a nanocrystalline matrix like those shown in Fig. l(a).

After the fibers have been alloy coated, they are consolidated to near net shape using either a hot isostatic pressing (HIP), vacuum hot pressing (VHP), or roll bonding approach. The ideal consolidation process results in complete densification [see Fig. l(b)] with neither the fibers nor their engineered interfacial coating incurring significant mechanical

1111 or chemical damage [12, 131. Failure to accomplish this results in fiber fractures and/or increased interfacial sliding resistance causing a loss

tTo whom all correspondence should be addressed.

of composite strength, creep rupture life, fracture toughness, and fatigue resistance [ 14, IS]. Models of the consolidated process have been developed for plasma spray deposited monotapes [l&18] and have been used for the design of optimal process schedules that limit damage during consolidation. The basic intent of models of this type is to analyse the micromechanisms of densification, fiber damage, and fiber-matrix reaction in a way that enables prediction of the relative density, fiber microbending/fracture. and the reaction layer thickness as a function of the process conditions. Inputs for the models are therefore the processing conditions (pressure and temperature as functions of time), the initial lay-up geometry, the matrix properties that govern densification, the fiber-matrix mechanical properties that control fiber damage mechanics and the Arrhenious parameters that determine interfacial reaction kinetics [16-181.

At first sight, the problems encountered in the densification modeling of metal coated fibers appear straightforward. The results of recent micromechani- cal analyses of contact deformations combined with the number of coated fiber-fiber contacts could be used to predict early stage densification [ 19, 201 while the analysis of circular cylindrical voids using appropriate potentials [21] would provide the required means of modeling the specimen as it approached full density. However, in reality the problem must also address the evolution of the number of contacts with density in a random packing of fibers, the effect of a rigid monofilament upon

1851

1852 KUNZE and WADLEY: THE DENSIFICATION OF METAL COATED FIBERS

Fig. I. Ti&6AlI4V sputter coated Sigma 1240 fibers. (a) Initial packing (as-received condition). (b) Fully dense sample after HIP consolidation (9OO’C, 20 MI% for 30 min followed by 100 MPa for 1 h).

contact deformation of the matrix, the potentially leads to an enhanced superplastic effect when the test non-circular shape of the pores created by coated temperature is within the CI + /J’ phase region (i.e. fiber-fiber contacts and the somewhat unusual matrix - 760&9OO”C) [lo]. However, the phenomenon is mechanical properties resulting from the nanocrys- complicated by the additional observation of rapid talline matrix produced by the PVD process. The grain coarsening accompanying exposure of the PVD latter issue arises because the PVD process is matrix to the elevated temperatures encountered conducted at relatively low temperatures where grain during the consolidation cycle. This could result in a growth is slow. Recent experimental studies with mechanical behavior that is a potentially strong similarly sputtered Ti-6A14V have shown that this function of the process path [lo].

Fig. 2 (a)

K

(a)

Uh IZE and WADLEY: THE DENSIFICATION OF METAL COATED FIBE

1.27

b-I

.RS

(b)

Test specimen

Eddy curre... sensor

knt 5.pa

(15124 1

Titanium tube (1.59 Dia. x 0.16 W)

/- Coated fibers

/ Evacuation groove (4)

, E-beam weld Titanium tapered end cap

Thermocouple center bore

Centering thermocouple

All dimensions in cm

rlindrical HIP canister assembly containing metal coated fibers. (b) Schematic 1 . easy current sensor

As a prelude to modeling the densification process, we have experimentally investigated the HIP consolidation process for silicon carbide monofilaments which had been sputter coated with a Ti-6A14V matrix similar to that used by Warren et al. [lo] in a detailed study of nanocrystalline creep. Using a hot isostatic press containing an eddy current sensor to monitor sample dimensions (and thus the sample’s density), the densification response of cylindrical specimens has been measured for five different process cycles. Several of the tests were deliberately interrupted to allow the evolution of the packing and the pore shape to be examined. The grain size has also been measured and compared to the predictions of a microstructural coarsening model for PVD Ti-6Al-4V [ lo].

diagl -am of

1853

2. EXPERIMENTAL PROCEDURE

2.1. Sample preparation

Tungsten cored Sigma 1240 fibers were used for the consolidation study. The Sigma 1240 fiber has a nominal diameter of 100 pm and a carbon-TiBz duplex coating [22]. The Ti-6Al-4V matrix was deposited on the fibers via a sputtering process at the 3M Company’s Metal Matrix Composite Center (Mendota Heights, MN) using similar conditions to those reported by Warren et al. [lo]. The near line-of-sight sputtering process resulted in an approximately elliptical matrix coating [Fig. l(a)]. During deposition the temperature was kept low (- 400°C) resulting in a - 30-100 nm matrix grain size [lo]. Upon cooling, the combination of an


asymmetric coating and the thermal expansion mismatch between the fiber and matrix caused a large number of the coated fibers to bend. The resulting curvature makes it difficult to produce a close packed fiber array and resulted in the random packing shown in Fig. l(a).

Bundles of approximately 4000 matrix coated fibers were packed in five cylindrical canisters [Fig. 2(a)] consisting of a tube and two endcaps of commercially pure (CP-2) titanium. The assembled canister had an overall length of 171 mm and a diameter of 15.9 mm. The wall thickness of the tube was 1.6 mm. The design of the canister allowed for a 51-mm long composite specimen. The effect of canister end constraint was reduced by the use of two endcaps that had conically tapered ends as shown in Fig. 2(a). With the taper, the tube wall was unsupported near the composite specimen and the applied load was directed onto the coated fibers. The circular cross-section of the endcaps also helped to reduce canister buckling and to maintain a circular cross-section in the composite specimen region. The endcap length was designed to position the coated fibers at the center of an eddy current sensor [Fig. 2(b)]. Four longitudinal grooves on the surface of each endcap were used to aid evacuation of the canister.

In order to avoid contamination of the matrix, both the tube and the endcaps were degreased, cleaned in a mild acidic solution, rinsed in water and ultrasonically washed in isopropyl alcohol. The tube and endcaps were allowed to air dry after which the coated fibers were positioned within the tube. After inserting the endcaps, the canister assembly was hermetically sealed by electron beam welding under

‘.O~ 0.9

0.6 -

0.4 - b l”“ld

4r 0.3 - d1n9, D = 1 .oo. ambient temperature

sensor reading. Dz 0.85. T= WYC

I .’ h&d censor readng. D = 0.699. ambient temperature 0.2

0.00 0.05 0.10 0.15 0.20 0.25 0.30

Real (Z/Z,)

Fig. 3. Normalized impedance plane diagram for eddy current sensor showing data corresponding to initial (I), intermediate (II), and final (III) stages of densification for experiment B (840% 20 MPa for 30min followed by

100 MPa for 1 h).

a vacuum of the order of 10 - ’ torr. Multiple measurements of the canister diameter were made prior to densification to determine the canister roundness and initial diameter. The maximum measured difference in the diameter for any of the specimens was 0.09% of the average initial diameter. The average of each measurement set was taken to be the canister diameter for the purpose of analysing the results and calculating the density.

2.2. Density determination

An eddy current sensor [Fig. 2(b)] installed within the HIP chamber was used to measure the canister diameter during each consolidation experiment. Details concerning the operating principle of the sensor have been described elsewhere [23-251. Briefly, the sensor used a two-coil technique in which a primary coil (when excited by a variable frequency current) induced an electromagnetic field of variable frequency that linked the electrically conductive specimen. This resulted in eddy current induction in the sample and a perturbation to the electromagnetic field. A secondary coil, located near the sensor surface [Fig. l(b)] sensed the field’s perturbation. Due to the electromagnetic skin effect, the field pertur- bations at high frequencies were functions of only the geometry of the component (and not its electrical conductivity) and this enabled the canister’s diameter to be deduced [23].

These effects were conveniently monitored by measuring the complex transfer impedance, Z, of the two-coil system with a general purpose multifre- quency impedance analyser (an HP 4194 network analyser was used in this study). In order to monitor accurately the change in specimen diameter during a consolidation cycle, the unperturbed electromagnetic field (i.e. the empty sensor impedance, ZO) must be known as a function of the process temperature. This was obtained by exposing the empty sensor to the same temperature profile used in the actual consolidation experiment and recording the sensor’s gain, G,, and phase &,, at each frequency sampled in the experiment. These readings were then used to normalize the gain, G, and phase, 4, data obtained during subsequent consolidation experiments. The normalized gain and phase measurements were subsequently used to compute a normalized sensor impedance, Z/ZO, at each test frequency using

Z - = $sin($ - 4”) + jcos(+ - &)I Z0 0 (1)

where Z, is the empty sensor impedance and Z is the impedance with the sample present.

For each experiment, the impedance was measured at 101 logarithmically spaced frequencies between 20 and 2000 kHz. The measurements were made at 1 min intervals with each frequency set taking approximately 23 s to complete. The data was then stored on a personal computer, along with the time of the reading. After the experiment was completed,

KUNZE and WADLEY: THE DENSIFICATION OF METAL COATED FIBERS 1855

Table I. Normalized impedance plane diagram slope and intercept values

Sensor reading Slope Intercept, h

I) Initial sensor readmg, D = 0.699 1.178 0.330 II) Intermediate sensor reading. D - 0.85 I.267 0.394 111) Final sensor reading, D = 1.00 1.186 0.466

the temperature/pressure data recorded by a computer monitoring the HIP was combined with the sensor data to obtain the densification response parametrically with each process variable’s (temperature and pressure) time dependence.

The first step in determining the density involved the analysis of the eddy current sensor output to calculate the specimen’s diameter. A normalized impedance plane diagram like that shown in Fig. 3 plots the imaginary and real impedance components as a function of test frequency. Figure 3 shows this “raw” data at three stages of one of the consolidation runs (experiment B). The data for the initial and final curves were taken at room temperature both before and after the experiment, while the intermediate curve was taken during the consolidation process when the relative density was approximately 0.85. The elevated temperature of this latter curve results in a reduction in the sample’s conductivity and thus a counter-clockwise shift in the lower frequency data (i.e. towards the empty sensor impedance). In spite of this, all three curves demonstrate a clear linear region at high frequencies that can be extrapolated to make an (infinite frequency) intercept with the imaginary axis. Values for the slopes and intercepts of these high frequency limits are given in Table 1 for each of the curves in Fig. 3.

An electromagnetic analysis of the sensor containing an infinitely long sample [23,25] has shown that the imaginary axis intercept of the high frequency (linear) portions of the normalized impedance curves corresponds to 1 - b, where h is the coil’s fill factor (i.e. the ratio of the sample’s cross-sectional area to that of the secondary coil). Thus, the apparent sample diameter, d:, can be obtained from

d,’ = d,&% (4

where d, is the diameter of the sensor’s secondary coil, and b is the high frequency intercept of the normalized impedance data.

It was found that the sample used here was not quite long enough for equation (2) to be strictly valid. The length of the secondary coil and composite sample were the same (51 mm) and the discontinuity in the deformation where the endcap and composite met induced a small additional perturbation in the electromagnetic field. This resulted in an apparent diameter that was slightly smaller than the actual diameter. Thus, the true diameter, d,, was related to the apparent diameter by

d, = d,‘[l + Ad,,,] (3)

where d: is the apparent diameter and Ad,,, is an end effect correction.

The small end correction was obtained by independently measuring (i.e. with a micrometer) the final diameters of the consolidated specimens and comparing this to the apparent value obtained by the sensor. It was found that Ad,,, was a function of the specimen’s total dimetral shrinkage, Ad,. Normalizing the data to account for the different final diameters, d(, yields an empirical relationship for the correction:

Once the high temperature diameter of the specimen was determined from the apparent diameter sensor reading using equations (3) and (4), the equivalent room temperature diameter was calculated from the thermal expansion relationship for titanium [26] and the room temperature density obtained.

2.3. Processing

An Asea Brown Boveri QIH-15 hot isostatic press equipped with a two-zone molybdenum furnace and four tungsten/rhenium thermocouples was used for this experimental study. Two of the thermocouples were used for the furnace control; the remaining two were used to monitor the specimen temperature. One thermocouple protruded from the base of the sensor [see Fig. 2(b)] and was also used to center the canister. The second thermocouple was placed in a bore hole located in the center of the top endcap. The HIP cycle was programmed via a personal computer which also logged the process cycle’s set points and vessel conditions (pressure, furnace temperature, and specimen temperature) throughout each experiment.

Each of the five cycles used in the consolidation study contained several similar segments (see Table 2). The first involved a chamber purge and pre-heat to remove moisture from the chamber. During this segment, the HIP chamber was evacuated, purged several times with argon, and re-evacuated. The furnace was then heated under argon to 300’C after which the chamber was evacuated a final time. The pressures and temperatures during this segment of the cycle were kept low to avoid densifying the specimen and affecting the microstructure of the matrix.

Once the chamber had been purged, the vessel was filled with argon to a pressure of 2 MPa (the minimum pressure required by the molybdenum

Table 2. HIP schedules

Experiment

Process variable A B C D E

Temperature ( C) 900 840 760 840 840 First pressure hold (MPa) 20 20 20 IO 5 Time at first pressure hold (min) 30 30 30 60 360 Second pressure hold (MPa) 100 100 100 N,‘A N,‘A Time at second pressure hold (mm) 60 60 60 NIA N*A


Table 3. Metallographic results of HIP specimens

Experiment

Prooertv A B C D E I ,

Initial density, D, 0.725 + 0.006 0.699 i 0.012 0.716 & 0.006 0.711 f 0.013 0.697 + 0.019 Final density, Dr 1.000 1 .a00 0.999 f 0.001 0.931 * 0.012 0.893 _+ 0.019 Volume fraction of fiber, L’, 0.449 f 0.014 0.494 * 0.015 0.483 * 0.014 0.468 i 0.010 0.468 + 0.008 Average number of contacts, 2 5.79 + 0.53 5.83 f 0.67 5.75 i 0.62 4.62 f 0.73 4.16 _+ 0.77

HIP furnace when operating at high temperature). The temperature was raised to the desired set point at a rate of lO”C/min. Once the temperature set point was reached, the system was held for 15 min to allow thermal equilibrium to be fully established within both the specimen and the eddy current sensor. Thermal equilibrium within the specimen was verified by monitoring the two thermocouples attached to each end. The furnace controller maintained a specimen temperature to within k 15°C of the set point temperature throughout the temperature hold.

After thermal equilibrium was attained, the vessel was pressurized to the first pressure set point. All five experiments used a relatively low initial pressure (relative to the 100 MPa typically used in HIPping conventional Tip6Al4V powders [27]) because of the unusually rapid consolidation associated with the enhanced superplastic nature of the matrix. This “gentle” HIP cycle has also been shown to minimize the fiber fractures that can occur as a result of fiber crossovers within a randomly packed fiber bundle [28]. Experiments A-C also used a second, high pressure hold to enable the monitoring of final void closure. The pressurization rate for the experiments was approximately 5 MPa/min when the chamber pressures were less than 7 MPa and about 1 MPa/min for pressures up to 100 MPa. After the pressure set point was achieved, the pressure was maintained to within kO.25 MPa of the target pressure.

The final segment of the cycle involved cooling the vessel and venting the argon gas. For temperatures above 400°C the cooling rate was 13”C/min. Below 400°C the cooling rate declined due to the lack of effective heat transfer between the HIP cooling system and the thermal mass of the furnace, sensor, and specimen. The argon was vented in stages simultaneously with the cooling, thereby aiding the cooling rate through the expansion of the gas within the chamber.

2.4. Post-consolidation characterization

After the samples were removed, their diameters were multiply measured to determine specimen roundness and verify the sensor readings. The maximum measured difference in any diameter pair was within 1.33% of the average final diameter. This indicated that the roundness of the specimen was well maintained during consolidation, validating an implicit assumption of a circular cross-section in the

eddy current computations of the specimen diameter [23, 241.

The specimens were then cross-sectioned via electrodischarge machining. The cross-sections were mounted in epoxy, polished and examined in a scanning electron microscope (SEM). Using an image analysis software package and at least 5-10 SEM micrographs (the greater number was used in analysing the partially consolidated specimens), the volume fraction of fiber, t+, and the final density, Df, were determined.

The measurements of the initial and final sample diameters (d, and df) together with the area of the canister wall, a, (assumed to be constant), and the final metallographically determined density Dr, enable one to calculate the initial density, D,, of the preform from conservation of mass:

The volume fraction of fiber is dependent upon the sputtering process (i.e. how much matrix is deposited). Although the sputtering process was designed to result in a predetermined fiber volume fraction, it is necessary to confirm the predicted value due to variations in conditions which can occur during the sputtering process.

2.5. Contact coordination number

The experiments can be used to determine the average number of coated fiber-fiber contacts, Z, as a function of the relative density. For the fully dense specimens [such as in Fig. l(b)], the number of contacts was determined by drawing Voronoi polygons [29] whose sides were located midway between adjacent fibers and were perpendicular to a line connecting each fiber center. Applying the approach Arzt [30] used for determining the coordination number in powder consolidation, each side of the polygon was taken to represent a contact. The average number of contacts for the partially consolidated specimens was determined from SEM micrographs by counting the number of contacts for each coated fiber and computing the average. The initial packing geometry was obtained by cross-sec- tioning an unconsolidated bundle of coated fibers which had been packed into a glass tube and impregnated with epoxy.


0.6

1000

900

600

700

600

500

-Region 1-j 2+-3-4-4-d I

i, :, ;r ;/ /# :, I-! ;I

:, ;r

:I ;:

-’ ,

Region l-

G--

2 MPa

_ ,I

0 100 200 300 400

Time (min)

500

Fig. 4. Time dependent data for experiments A, B, and C. (a) Density, (b) temperature. and (c) pressure histories.

3. RESULTS AND DISCUSSION

3.1. Metallography results

The results of the metallographic analysis for each of the five experiments are summarized in Table 3. The uncertainty in the computed values for the initial density arises from the variability in the values of the final density and the initial and final diameters. Despite this, the initial relative density varied less than 3% with an average of about 0.71. The slight

difference in the starting densities may be due to the random packing of the bundle and to a small uncertainty in the values used in equation (5). Metallographic analysis of each composite revealed that the fiber volume fraction averaged about 47% and varied less than 5%. The relatively constant volume fraction of fiber means that the effect of the fiber on the consolidation process should be nearly the same for all of the experiments.

3.2. Densification kinetics

By letting the initial density, D,, and diameter, dIt,, in equation (5) be the sample density and diameter at any time and using equations (2), (3) and (4), it was possible to determine the densification behavior of each of the specimens from the data taken by the sensor throughout the cycle. The first three specimens, A-C. were consolidated at temperatures of 900’. 840”, and 76O’C. The same pressure profile (20 MPa for 30 min and then increasing to 100 MPa and holding for 1 h) was used for each of the three temperatures (see Table 2). The specimen temperature (taken as the average of the two attached thermocouple readings), chamber pressure, and the computed relative density for experiments A, B, and C are shown in Fig. 4. The density at the various regions depicted in Fig. 4 are given in Table 4.

In all three experiments, a chamber pressure of 2 MPa was used while the temperature was raised to its predetermined set point. For experiment A, where the set point temperature was the highest (900°C) even this small pressure increased the density to 0.794 before the first pressurization ramp began [Fig. 4(a)]. A lesser effect was seen for experiment B (840 C) and experiment C (76O’C). There are two possible explanations for this behavior. One is a decrease in the matrices’ resistance to flow as the temperature is increased leading to densification at very low pressures. The second may be a result of a constraining (i.e. load supporting) effect of the canister at the lower temperatures. Titanium under- goes an cc-p phase transformation at 882°C which is between the test temperatures of experiments A and B (i.e. the 900 and 84O‘C experiments). A transition from the P-phase to the more creep resistant c1 phase causes the creep rate to decrease by about a factor of 10 for the stresses applied here [31]. Decreasing the temperature further from 840 to 760°C has been

Table 4. Densities at various points during HIP consohdation

Experment

Density A B C D E

htial density, D, Start of region I First pressurization density, DpI start of region 2 Second pressurization density, Dpz Start of region 3 Final density. DI End of region 4

0.725 0.699 0.716 0.711 0.697

0.794 0.720 0.72 I 0.738 0.720

0.980 0.956 0.908 WA N,‘A

I.000 I .ooo 0.999 0931 0.893


25 I I I I I I

fc)

c 20 - :'--' 2 B

15- ;

P : lo- I E 5-

0 lUl

---____ I I I I I

0 100 200 300 400 500 600 700

Time (min)

Fig. 5. Time dependent data for low pressure experiments B, D, and E. (a) Density, (b) temperature, and (c) pressure

histories.

found to result in an additional decrease in the creep rate by about a factor of 7 [31].

When the pressure was raised to the first pressure hold of 20 MPa, all three specimens showed a very rapid increase in density. It is evident from Fig. 4 that the densification rate in this regime is strongly temperature dependent. Again, experiment A densified most rapidly and reached a relative density of 0.98 by the end of the first pressure hold (end of region 2 in Fig. 4). Experiment B showed a slightly less rapid increase and reached a relative density of about 0.95 before the second pressurization ramp began. Experiment C exhibited the lowest densification rate of the three but still attained a relative density of about 0.90 by the end of region 2. Decreasing the temperature resulted in both a decrease in the densification rate and the density at which the material displayed a “hardening” effect. However, all three experiments reached a relatively high density in a short amount of time and attained a relative density of about 99% by the time the high pressure set point of 100 MPa was reached.

Both experiments A and B reached full density. Experiment C reached a final relative density of 99.9%. Final closure of the remaining voids would have required a higher temperature or pressure, or possibly a longer hold time. These three experiments indicate that these metal coated fibers can be fully densified at 100 MPa at temperatures well below the 9OO.C normally used for conventional Ti-6Al4V powders [27].

Two additional lower pressure experiments (D and E) were conducted at 84OC and deliberately

Fig. 6. Partially consolidated specimen D showing a distribution of void sizes and shapes (840°C and 10 MPa for 1 h).


Fig. 7. (a) Example of an early stage void in specimen E (840°C and 5 MPa for 6 h) and (b) a large Type I void in specimen D (840°C and 10 MPa for 1 h). These correspond to regions of low initial packing and reveal the way a non-uniform coating and the random fiber packing result in large multicornered voids

created by multiple fiber contacts.

interrupted before full density was reached. Their relative densities, temperature, and pressure histories are shown in Fig. 5 along with the low pressure segment of experiment B which was also conducted at 840°C. Taken together, the three experiments show surprisingly high densification rates (given the low pressures and relatively moderate temperature used). The results indicate that significant creep densification occurred during

the pressure hold in all three cases. The rate densification decreased with decreasing presst Experiment B reached a density of 0.956 at the c of the first pressure hold, experiment D reacl a final density of 0.931, and experiment E reacl a final density of 0.893. In al1 three cases, appeared that further densification would h. occurred if the experiments had been allowed continue.

of u-e. 2nd led led

it ave

to


3.3. PI ore shape

Met allographic analysis of specimens D and E indicat ed that they contained significant porosity. A typical cross-section is shown in Fig. 6. It reveals a signific :ant variability in local relative density. For instanc :e, a high relative density region is evident near the tof I of the figure while a significantly lower density region exists near the center. This variability in relative : density is presumably a consequence of the

initial random fiber packing of the specimen results in regions of locally low initial packing d [Fig. l(a)].

which ensity

The samples contained no spherical voids. R ather, a distribution of cusped pores existed with a rai lge of sizes and shapes. Two main types of pore were present (Fig. 6): Type I consisted of large cusp si haped voids whose sides are composed of several fiber coatings, usually four. Type II were typically sn Taller, usually had three corners, and appeared to have

Fig. 8. High magnification views of Type I (a) and Type II (b and c) voids in specimen D (840°C and 10 MPa for 1 h).


Fig. 8(c)

resulted from the pinching-off of the larger Type I pores. The micrographs in Fig. 7 are typical Type I voids in regions of locally low relative density. Figure 7(a) depicts a region from specimen E where the local relative density is low and fiber contacts are growing to form Type I and Type II voids. In this example, two contacts are about to occur near the center of the photo that will lead to the formation of three-sided “triangular” Type II voids in the center and left while a four-sided “rectangular” Type I void forms on the right. The large five-sided void shown in Fig. 7(b) is an example of variability in void shape resulting from the random fiber packing. If a coated fiber had been present where the center of the void now exists (i.e. if a locally close packed structure had existed), the area would have been a highly dense region with small Type II voids. Since the matrix must flow a considerable distance, one might expect that this type of void will be difficult to densify.

Figure 8 shows the formation of Type II voids from the larger Type I by a pinch-off mechanism. The pore shown in Fig. S(a) is a typical four-sided Type I void in a region of moderate local relative density. The two triangular voids shown in Fig. 8(b) appear to have formed by the pinching off of a void like that in

Table 5. Material properties of Ti-6A14V B-phase [IO]

Material parameter P-phase

Atomic volume, R (m’) 1.7 x lo-‘? Grain boundary width, d (m) 6 x lo-‘” Grain boundary energy, r (J/ml) 0.35 Grain boundary diffusion pre-exponential factor,

DOPh (n?/s) 1.3 x lo-’ Grain boundary chffusmn activation energy,

Q~:D (J) 1.28 x lo-

Fig. S(a). Figure 8(c) shows a smaller Type II void from a region of high local relative density. Note that the corners of this void have remained sharp. The smaller Type II voids remained cusp shaped even when exposed to high temperatures for over 6 h (experiment E). This must mean that the creep deformation rate significantly exceeded the rate of pore rounding via surface diffusion [32] (i.e. the deformation rate associated with enhanced superplas- ticity exceeded the diffusional sintering rate). This observation has considerable practical significance since the densification rates of cusp shaped voids have been shown to densify 2-3 times faster than spherical voids of similar volume fraction [33, 341.

7- ru m 6- 5

0

2.9 5-

01 I / I I I

0.7 0.8 0.9 1 .o

Relative density D

Fig. 9. Measured variation in average number of coated fiber-fiber contacts as a function of relative density during

HIPping.


Table 6. Grain growth parameters for PVD Tim6A1-4V [lo]

Temperature do k a

900 0.50 0.23 0.20 840 0.20 0.23 0.20 760 0.11 0.14 0.24

Early models for the densification of powders [35-371 and more recent models for plasma sprayed composite monotapes [l6] have all assumed that the voids present during final stage consolidation were spherical (the equilibrium shape). This assumption implies that surface diffusion along the void surface is sufficiently fast to continuously redistribute material being deposited from the contact by creep/boundary diffusion. Recently, pore shape models have been developed for quasi- and non-equilibrium shaped pores [38, 391. From the results of Svoboda and Riedel [39], who modeled the sintering of a hexagonal array of wires, it is possible to estimate the time required for the pores to reach a near spheroidal (quasi-equilibrium) shape. Since Ti-6A1-4V is a duplex alloy, a conservative estimate can be made using the data for /3-Ti given in Table 5 (/I-Ti has a higher diffusivity than cc-Ti). The radius of the coated fiber can be estimated for a uniformly coated fiber with a fiber volume fraction of 47%. Such a coated fiber has a radius of approximately 73 pm. The Svoboda and Riedel model then predicts that a quasi-equilibrium shape will be attained after a time of the order of IO6 h at 900°C. This result arises from the relatively low diffusivity of /I-Ti (compared to other metals), and the large radius of the coated jber. If the coated fiber radius were one-tenth the size (7.3 pm), a quasi-equilibrium shape would be achieved of the order of lo2 h. The larger radius of the coated fiber results in larger diffusion distances and in a significantly longer time for the pores to reach a quasi-equilibrium shape.

3.4. Contact coordination number

The evolution of contact coordination number with relative density provides a mechanism for mapping the calculated densification of individual contacts to the ensemble body. A metallographic analysis has enabled the number of coated fiber-fiber contacts to be measured as a function of density. The average numbers of contacts for all the consolidated specimens are given in Table 3. The average coordination number for the unconsolidated bundle was found to be 2.3 f 0.99 at a relative density of

Table 7. A comparison of measured and predicted grain sires after HIPping

Measured grain size Predicted grain size Experiment (rem) (pm)

A (900 C. - I90 min) 1.99 f 0.15 1.99 B (840 C, - 190 min) 1.76 k 0.09 1.69 C (760 C, - I90 min) I .45 * 0.04 I .43 D (840’ C, z 75 min) I .37 5 0.12 1.44 E (840°C. - 380 min) I .89 i 0.07 1.91

Applied pressure (MPa)

5 10 15 20 I I I 1

Ti-6AL4V/Sigma 1240 T= 840% (TI T,,,= 0.576)

0.95

.z 0.90

E 8 a, 0.85 .z m 5 = 0.80

I I I 0.1 0.2 0.3

Normalized pressure P,, /oy

- 3.0

- 2.8

- 3.6

- 3.4 B

.*

3.2 E - 0”

Fig. 10. Experimental HIP map for 840°C depicting relationship between relative density, pressure, and time.

0.70, nearly identical to the initial density of the consolidated specimens. The data is plotted as a function of relative density in Fig. 9. The results form a linear relationship between the average number of contacts and density:

z = 10.80 - 5.3. (6) At full density the number of contacts was close to the number of contacts for a hexagonal close packed array, i.e. 6. From the work of Arzt and others [30, 351 on the consolidation of powders, the number of contacts of a powder particle, Z,, as a function of the relative density of the powder compact, D,, may be expressed as Z, = 120, [35]. For powders, the number of contacts at full density is 12 and arises from the three-dimensional nature of a powder packing compared to the two-dimensional case of aligned coated fibers.

The curvature of the coated fibers due to the thermal expansion mismatch between the fiber/ matrix and the subsequent cooling at the end of the deposition process results in a random packing of the

Temperature (“C)

700 750 800 850 900 950 1.00, I I I I , ,

Ti-6Al-QV/Sigma 1240

0.50 0.52 0.54 0.56 0.56 0.60 0.62 0.64

Homologous temperature T/T,

Fig. 11. Experimental HIP map for 20 MPa depicting relationship between relative density, temperature, and time.


coated fibers with an average initial relative density of -0.70 [Fig. l(a)]. Regions of higher and lower density also existed and persisted during subsequent densification (Fig. 6). The overall lower packing density of the random bundle results in an initial average of 2.3 contacts per coated fiber. This means that in the early stages of densification each contact must carry a larger portion of the load (and hence support a higher contact stress) than in a hexagonal close packing of fibers subjected to the same pressure. This can contribute to the observed high densification rates in Figs 4 and 5. As samples densified, both the average contact area and the average number of contacts increased [according to equation (6)]. Thus the average contact stress would have dropped for a constant applied pressure, resulting in the rapid decreases in the densification rate observed during the pressure holds.

3.5. Grain growth kinetics

At elevated temperatures the small (- 50 nm) grain size of the PVD deposited Ti-6AlAV matrix rapidly coarsens. In the study of the deformation of PVD Ti-6Al-4V, Warren et al. [lo] found the grain growth kinetics to be well represented by an empirical expression of the form

d = d,, + kt” (7)

where d is the grain size (in microns) after isothermal thermal exposure for a time t (in seconds) and d,, k, and a are temperature dependent parameters given in Table 6 for each of the temperatures used in the Warren et al. [lo] study.

The matrix grain sizes were measured for each of the specimens consolidated in this study using Hilliard’s single-circle intercept method [40]. The grain measurements shown in Table 7 reveal that as the consolidation temperature was raised, the final grain size increased. After an initially rapid growth from the submicron range, further grain growth became retarded. This may be seen by comparing the final grain sizes of experiments D, B, and E which were conducted at 84O’C but for increasing amounts of time.

The predictions of equation (7) using the constants in Table 6 are also given in Table 7. The close agreement between these two sets of results indicates that neither the presence of the fiber, the annular geometry of the fiber coating, nor the strain accompanying consolidation significantly affect the relationships developed by Warren rt al. for the grain growth kinetics of PVD sputtered Ti-6Al4V foil DOI.

It is interesting to note that after consolidation the measured matrix grain sizes were between 1.4 and 2 pm where as conventionally processed Ti-6A14V has a typical grain size of 3-7 pm after a similar treatment [41]. Warren et al. [IO] have shown that between -740 and 900 C, the matrix creep is well represented by the Ashby-Verral model [42] for

superplastic flow in which the strain rates are proportional to I/&. Thus, the PVD matrix, even at the end of consolidation, is likely to creep as much as an order of magnitude faster than conventionally processed Ti-6A14V alloy exposed to the same stress. Even higher rates may be possible during the early stages of consolidation when the grain size is even smaller. For example, from equation (7) the grain size after 2 min at 84O’C would be 0.8 pm and so the creep rate would be an additional four or five times faster than at the end of the test. This rapid creep rate during the early stages of consolidation is likely to have significantly contributed to the anomalously rapid densification observed in these experiments. The combination of rapid grain growth kinetics and the I/& creep rate dependence on grain size indicate that it will be important to introduce this coarsening process into future densification models for PVD coated materials.

4. EXPERIMENTAL HIP MAPS

The density of a porous material undergoing consolidation is a complicated function of the pressure, temperature and time. Ashby and co- workers [35,37] have developed HIP maps as a convenient means of representing the effect of HIP processing conditions (temperature, pressure, and time) on density. HIP maps show constant time plots of the relative density as a function of one of the process variables (temperature or pressure), while holding the other constant. Usually the isochronal density relationships are obtained from densification model predictions [35, 371. HIP maps of this type have been developed for the consolidation of plasma sprayed metal matrix composite monotapes [16] and the densification of monolithic alloy powders [36, 37,431. They are not only helpful in identifying the best conditions for processing, but also in identifying regions where single deformation mechan- isms dominate densification.

HIP maps can also be constructed from experimental data and provide a means for condensing data into a format that can be conveniently compared with model predictions. Choi et al. [43] constructed experimental HIP maps of this type for the consolidation of titanium aluminide powders using data from more than 40 interrupted HIP experiments. Here, we are able to construct a HIP map using the eddy current measured densification data from the three experiments conducted at 840’ C. The resulting map (Fig. 10) shows the relative density as a function of pressure (normalized by 54 MPa, the yield strength of Ti-6A14V alloys at 840,‘C [16]). Five time contours ranging from 1 to 30 min are shown. All start at the same initial density (71%) which was the average of the starting densities. Time was measured from the point at which the pressure reached its low pressure set point. From the map, it is clearly evident that as the density is initially increased, the pressure


required for further densification rises almost linearly. A flattening of the isochronal curves at higher densities is a result of the increase in contact area and number of contacts (thus decreasing the contact stress) as the density increases.

A HIP map showing relative density as a function of temperature at 20 MPa is shown in Fig. 11. In this case, the map shows a nearly linear increase in each isochronal density curve with temperature. However, the slopes of these relationships decrease with time. This results in a convergence of the isochronal contours as the temperature increases. The overall rise in density with temperature results from an increase in the matrix creep rate with temperature [lo]. The more rapid “consolidation time” hardening (i.e. the convergence of the isochronal curves) at high temperature may be a manifestation of the more rapid grain growth (and thus creep hardening) superim- posed on the hardening due to the increase in contact area and number of contacts. The maps shown in Figs 10 and 11 reveal that relatively low pressures (lo-20 MPa) at moderate temperatures (850-900°C) result in very significant densification.

5. SUMMARY

1. Cylindrical specimens containing randomly packed SIC monofilaments sputter coated with Ti-6A14V have been HIP consolidated and the densification response measured using an eddy current sensor. The sensor was used to systematically determine the effect of the process temperature, pressure, and time upon densification. A significant increase in density occurred during initial temperature ramping, even when the pressure was low (N 2 MPa). The densification rate was a strong increasing function of pressure and temperature. 2. Experimental HIP maps have been constructed from the experimental data. They reveal that these materials are rapidly densified at surprisingly low pressures. The maps conveniently summarize the effect of varying the process variables upon the densification behavior and enable one to determine the conditions needed to successfully densify matrix coated fibers. 3. A study of the evolution of the internal geometry (coated fiber contacts and void size, shape, and distribution) during consolidation has revealed that initially, the average number of contacts per coated fiber was about 2 and was a linear function of the relative density. The pores present during final stage densification were cusp shaped even after consolidation at the highest temperature. This arises from the large surface diffusion distance encountered in the coated monofilament geometry. 4. Rapid grain growth accompanied consolidation of the initially nanocrystalline matrix. An empirical model for grain growth kinetics of monolithic PVD Ti-6A14V developed by Warren et al. has been found to accurately describe the grain growth kinetics of the

fiber reinforced matrix. The refined microstructure of the matrix resulting from the PVD process may have enhanced the superplastic behavior of this alloy and allowed the metal coated fibers to be consolidated at pressures and temperatures well below those required for the HIPping of conventional Ti-6Al4V powders. Models that seek to predict and optimize the consolidation process will need to incorporate this microstructure behavior.

Acknowledgements-We are grateful to the Defense Advanced Research Projects Agency (W. Barker, Program Manager) and the National Aeronautics and Space Administration (D. Barker, Technical Program Monitor) who supported this work through NASA Grant NAGW 1692.

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