PROCESSING EFFECTS ON THE HIGH CYCLE FATIGUE
LIFE OF WELD REPAIRED CAST Ti-6A1-4V PARTS
by
GORDON BRUCE HUNTER
SUBMITTED TO THE DEPARTMENT OF MATERIALS SCIENCE AND
ENGINEERING IN PARTIAL FULFILLMENT OF THE REQUIREMENTS
FOR THE DEGREES OF
MASTER OF SCIENCE
and
BACHELOR OF SCIENCE
at the
MASSACHUSETTS INSTITUTE OF TECHNOLOGY
June
©
1981
Gordon Bruce Hunter, 1981
The author hereby grants to M.I.T., the U.S. Army and any
agency of the U.S. Government permission to reproduce and
to distribute copies of this thesis document in whole or
in part.
7
Signature of Author
Department of Materials Science and Engineering
/
M;;v 8, 1981
Certified by
I Thomas W. Eagar
iThesis Supervisor
Accepted by
I
Regis M.N. Pelloux
J hhairman, DepartmentkGraduate
AR1CHIVES
a
MASSACHNLUSETTSST
JUL 1 7 1981
L1BRARIES
ira,
Committee
John B. Vander Sande
Department Undergraduate Committee
-2PROCESSING EFFECTS ON THE HIGH CYCLE FATIGUE
LIFE OF WELD REPAIRED CAST Ti-6A1-4V PARTS
by
Gordon Bruce Hunter
Submitted to the Department of Materials Science and Engineering
on May 8, 1981 in partial fulfillment of the requirements for the
Degrees of Master of Science and Bachelor of Science at the
Massachusetts Institute of Technology.
ABSTRACT
Specimens were processed according to a twelve-run Plackett-Burman
statistical screening design to study the effects of weld repair on high
cycle fatigue properties of cast Ti-6A1-4V. Variables studied were betatreatment before welding, weld bead size, post-weld hot isostatic press
treatment, heat treatment time and temperature, heat treatment cooling
rate, surface polish, and peening. Specimens were tested by R.R. Moore
rotating beam fatigue machines at 72.5 ksi and 10,000 rpm, and then examined by metallography,fractography, x-ray radiography, RMS surface
measurements, and chemical analysis. The results were also compared to
the high cycle fatigue properties of unwelded specimens.
The analysis of the results showed that weld repair is not detrimental to high cycle fatigue properties of cast Ti-6A1-4V. Surface condition was shown to be the most critical variable with a finer surface
finish and a higher surface compressive stress improving fatigue properties. Mechanical polishing was superior to glass bead peening, which
was superior to chemical milling. Microstructure was also shown to be a
critical variable with a larger colony size improving fatigue life by
slowing fatigue crack growth. The weld zone resisted fatigue crack
initiation due to a basketweave structure and thin grain boundary alpha.
A secondary hot isostatic press (HIP) treatment was shown to be unnecessary if the weld zone is sound. A HIP treatment may be made either
before or after welding without detrimental effects. Multiple treatments were shown not to be detrimental.
Multiple pass repair welds did not affect fatigue life when compared to single pass welds.
Thesis Supervisor:
Title:
Dr. Thomas W. Eagar
Associate Professor of Materials Engineering
- 3-
TABLE OF CONTENTS
PAGE
TITLE PAGE
1
ABSTRACT
2
TABLE OF CONTENTS
3
LIST OF FIGURES
5
LIST OF TABLES
7
ACKOWLEDGEMENTS
8
CHAPTERS
I.
II.
III.
IV.
V.
VI.
VII.
INTRODUCTION
10
PREVIOUS WORK
19
EXPERIMENTAL PROCEDURE
24
A.
Statistical Design
24
B.
Prior Microstructure - Variable 1
28
C.
Bead Size - Variable 2
33
D.
Hot Isostatic Press - Variable 3
35
E.
Heat Treatment - Variables 4 and 5
35
F.
Surface Finish - Variables 6 and 7
38
G.
Testing Methods
40
RESULTS
42
DISCUSSION
65
SUMMARY AND CONCLUSIONS
70
POSSIBILITIES FOR FUTURE WORK
72
-4APPENDIX I.
EFFECT OF TRIAL 5 RESULTS
73
APPENDIX II. FAILURE STRESS AT FRACTURE
76
BIBLIOGRAPHY
79
-5LIST OF FIGURES
FIGURE
PAGE
1
Solar prototype APU impeller in the as-cast condition.
16
2
Early DDA prototypes for a cast impeller. Impeller on
right is as-cast and impeller on left has been chemically
milled to final contours.
17
One of first two
of mold.
18
3
UH60A main hub casting, as broken out
4
Diagram of tube furnace.
32
5
Diagram of a machined blank.
34
6
Diagram showing positions of weld passes.
37
7
Diagram of the dimensions for the specimens after
machining.
39
8
Specimen 4 base metal.
48
9
Specimen 19 base metal.
48
10
Specimen 36 base metal.
49
11
Specimen 38 base metal.
49
12
Specimen 1 weld metal.
50
13
Specimen 32 weld metal.
is shown at bottom.
Fracture through the weld met,al
50
14
Specimen 37 weld metal.
51
15
Specimen 52 weld metal.
51
16
Specimen 1 microstructure showing transition from weld
metal to base metal. Weld metal is shown in the upper
part of the figure. The bottom of the figure shows the
fracture plane across base metal.
52
Specimen 10 fatigue origin showing fracture facets.
Lower part of figure is the outer surface of the
specimen.
57
17
-6-
18
Specimen 9 fatigue origin showing fracture facets. Lower
part of figure is the outer surface of the specimen.
57
19
Specimen 13 fatigue origin in weld metal.
58
20
Specimen 21 fatigue origin showing smeared appearance.
58
21
Specimen 7 outer surface showing mechanically polished
condition.
59
Specimen 3 outer surface showing chemically milled
condition.
59
Specimen 10 outer surface showing mechanically polished
and peened condition.
60
Specimen 4 outer surface showing chemically milled and
peened condition.
60
22
23
24
25
Specimen 9 outer surface showing peened condition.
(a) shows the secondary electron image with the fracture
origin and part of the fracture surface in the upper
right section of figure. (b) is an x-ray map showing
concentrations of silicon. Note that the dark areas in
(a) correspond
to silicon
conditions
in (b).
61
26
S/N plot
27
S/N plot for welded and unwelded specimens.
64
28
Plot showing the effect of various values for trial 5
on the statistical results.
75
Diagram showing the method of determining the specimen
radius at the fracture.
77
29
for the screening experiment specimens.
63
-7LIST OF TABLES
PAGE
TABLE NUMBER
1
Twelve-Run Plackett-Burman Design
25
2
"t"
Distribution
27
3a
Rod Composition
29
3b
Rod Mechanical
4a
Process Variables
30
4b
Rod Processing
30
5
Welding
6a
R.R. Moore
Results
43
6b
Additional R.R. Moore Results
43
7a
Fatigue Life Statistics
44
7b
Fatigue Life Significant Effects
44
8
Specimen Surface Finishes
46
9
Slug Analysis
47
10
Specimen Colony Size
54
lla
Colony Size Statistics
55
llb
Colony Size Significant Effects
55
12
Comparison of Fatigue Life and Colony Size
56
13a
Effect
of Trial
5 on HCF Factor
74
13b
Effect
of Trial
5 on HCF Significant
14
Calculated Failure Stresses
Properties
29
Schedule
Parameters
36
Effects
Effects
74
78
-8ACKNOWLEDGEMENTS
Most of the work contained in this report was supported by AMMRC in
Watertown, MA through the cooperative studies program of the Department of
Materials Science and Engineering at MIT.
Additional funding for work in
MIT's weld lab was provided by Mr. Bruce MacDonald of the Office of Naval
Research.
The author is grateful for the support of all three agencies.
Special thanks go to all the guys in the MIT welding lab, especially
Mr. Bruce Russell,and everyone in the Prototype Development Division at
AMMRC.
Everyone has been very kind and more than helpful in the past
year.
The author would also like to acknowledge the expert and friendly
help of Mr. Andy Zani and Ms. Carolyn Shimansky at AMMRC with the metallography and Mr. Andy Connolly of AMMRC and Ms. Libby Shaw of MIT with the
electron microscopy.
Ms. Rosalie Allen has been a godsend for her speedy
and accurate typing job, as well.
Mr. Al Hammer of Solar and Mr. Stan Silverstein of Sikorsky have
been very helpful with technical advice.
Professor Thomas King of MIT
has also been of great service by finding this co-op assignment for me.
It would
be negligence
to fail
to mention
the friendship
and
support of all the friends, associates, "family" and the UGOI who helped
maintain my sanity and the spirit of M.O.S., Inc.
The author is indebted to the advice, help and friendship of both
Mr. Frank Hodi and Dr. Robert French of AMMRC.
Nothing I could ever write expresses my gratitude to Professor
-9-
Tom Eagar.
He continues to be a great teacher, employer, advisor, and,
above all, friend.
Ms. Linda Schaffir has been a wonderfully stablizing influence in
the massive confusion of a college education.
Besides aiding the author
with the figures in this paper, she has been the best companion a guy
could ever know.
And, of course, there is Mom and Dad.
They stood by me when I
couldn't keep my "d"'s and "b"'s straight and when I couldn't ski down
the big hill, having to walk instead.
No matter how cliche it might be
(and this author hates cliche), this is dedicated to them.
- 10 I.
INTRODUCTION
In an effort to reduce the cost and lead time for the manufacture
of rotating titanium aircraft components, various commercial contractors
with support from the U.S. Army
have been studying possibilities for re-
placement of Ti-6 wt% A1-4 wt% V (Ti-6-4) forgings with titanium castings.
[1-4]
These parts can be cast to near net shape, reducing the amount of
scrap metal and reducing or avoiding time-consuming and expensive machining operations.
While most mechanical properties of the cast material
have been comparable to the forged material, the fatigue test results
of the castings have generally been inferior and have been characterized
by high scatter. [5-8]
Weld repair to ensure the integrity of cast com-
ponents is necessary to reduce the rejection rate, thus making castings
more attractive economically when compared to forgings.
There has been
concern that welding would degrade the fatigue properties and make castings unacceptable for rotating components.
In general, Ti-6-4 castings have a tendency for two types of internal defects; inclusions near the mold interface, and voids due to gas
porosity and solidification shrinkage.
[5, 6, 9, 10]
Both types of
defects have been shown to reduce fatigue strength. [7]
The inclusion
problem, in many cases, has been solved by casting the components slightly
oversize and then etching the surfaces away by chemical milling.
The
milling is continued until final dimensions are attained, eliminating
surface inclusions and any casting investment remaining.
Chemical mill-
ing is also used to remove contaminated surfaces formed during heat
- 11 treating.
Hot isostatic press (HIP) treatments have been successfully
used to eliminate internal voids. [6, 8, 9, 11]
Further heat treatments
are generally needed to improve the microstructure, increase the fatigue
strength, and reduce the scatter in properties.
Closure of internal voids during the HIP cycle may produce dimples
on the surface of the component.
Further, there can be inclusions in the
structure which are too deep to be removed by chemical milling and must
be machined out.
In either of these cases, weld repair is employed to
backfill and build-up material.
The most popular method of weld repair
employs gas-tungsten arc welding (GTAW) with argon gas and Ti-6-4 ELI
(extra low interstitial content) filler wire.
[6, 11, 12]
Following
welding, the component is either stress relieved, or otherwise heat
treated to prevent distortion in service.
follows
welding.
A second HIP treatment often
[6, 9 10, 13]
Repair welding is generally accomplished by making repeated
passes, laying down filler metal so that the part will meet specified dimensions.[12]
Small surface defects may be closed by a single shallow
pass without filler metal.
[12]
High purity must be maintained during
welding as small amounts of oxygen may affect the mechanical and fatigue
properties of the material.
[14]
Following weld repair, a HIP treat-
ment is sometimes required to ensure closure of any porosity or voids
that may have formed in the weld zone.
No published studies show the
effect of bead size or number of weld passes on internal stresses, microstructure, or mechanical and fatigue properties of cast Ti-6-4.
Few, if any, of the applications which are anticipated would re-
- 12 -
quire components with the as-cast surface intact because of the variability of the as-cast finish and because of processing procedures, such as
weld repair, mechanical machining, and chemical milling.
Additionally,
chemical milling is required to remove the surface casing if the heat
treatments are conducted in a contaminated atmosphere.
Many components
are glass bead peened for a final surface finish, inducing surface compressive stresses which improve the fatigue properties. [15, 16]
Titanium exists in two allotropic forms: a lower temperature closepacked hexagonal structure (),
and a body-centered cubic structure ().
Because of the relative stabilizing effects of the alloying elements,
Ti-6-4 has a mixed a +
structure with the
-transus near 1823°F (995°C).
[17]
Cast material cooled from the
structures.
The basic morphology is a Widmanstatten structure with
broad a-platelets separated by
[18]
-phase region has several common
-films and a complex interfacial phase.
The platelets are arranged in either a
structure. [19]
A colony is a packet of similarly aligned and crystallo-
graphically oriented platelets. [19, 20]
the prior
basketweave or colony
Long a-platelets also form in
-grain (PBG) boundries, called grain boundary (GBa). [21]
Additional heat treatments can form martensitic structures (a' and a"),
or recrystallized primary a-grains interspersed in the Widmanstatten
structure.
22]
The morphology, colony size, and a-platelet width depend
on the cooling rate, the concentration of
B-solutionizing conditions. [18-20, 22]
-stabilizing elements, and the
Rapid quenching will aid the
formation of a martensitic structure in preference to a Widmanstatten
- 13 -
structure. [22]
or with increased
Colony size decreases with an increase in cooling rate
-stabilizer content. [20]
A basketweave structure re-
sults when the colony size is so small that the transformed region appears
to be a distribution of a-platelets in random orientations.
lets and wide GBa are formed during slow cooling. [20]
Wide plate-
Heat treating
titanium in the presence of oxygen causes an a-casing to form due to the
a-stabilizing character of oxygen.
This casing usually does not excede
about .010" (.25 mm) in depth. [9]
Many published studies have shown that PBG size and colony size
have an effect on fatigue crack initiation and growth in a +
alloys.
8,
19, 20, 23-29]
titanium
The similar crystallographic orientation of
a-platelets causes slip and shear across a colony in certain favorable
orientations, thus facilitating fatigue crack initiation. [19, 23]
While
colonies promote crack initiation, they also promote branching and deviation of the main crack from its preferred course by the same slip
mechanism thus reducing the crack growth rate, as shown by Yoder, Cooley,
and Crooker.
23-27]
As fatigue cracks must reinitiate at each colony
boundary to propogate through the material, Eylon and Bania concluded that
a smaller colony size would slow crack growth.[20]
Based on these results,
Yoder and Eylon suggested that the fatigue crack growth rate decreased with
increasing colony size if the specimen thickness was much larger than the
colony size and that the rate increased with increasing colony size if
the colony size approached the thickness of the specimen. [28]
This
means that for a given specimen thickness, there is a colony size that
corresponds to a minimum growth rate.
- 14 The effective colony size is increased if two or more colonies
share a weak a-basal plane and can thus shear together. [29]
Because
the orientation of colonies is related to the parent PBG orientation by
Burgers relations, the largest effective colony size would theoretically
be limited by the PBG size. [29-31]
Eylon concluded from this that a
larger PBG size would tend to increase the effective colony size and
therefore reduce the fatigue crack growth rate in specimens with small
colony
size.
[28, 29]
Under certain conditions, GBa can promote fatigue crack initiation.
Eylon
found that subsurface GBa is a favorable initiation site in Ti-6-4,
especially if the a-platelet is planar, is inclined at or near 450 to the
tensile axis and is parallel to the long axis of a-platelets in an adjacent colony.[8,2l]Larger PBG size increases the number of large inclined
planar GBa platelets.
The coarse microstructure and variation of structures, cause the
wide scatter bands in fatigue results of cast Ti-6-4. [6, 20]
If the
component is to perform reliably, the low end of the scatter band must
meet the minimum fatigue specifications.
scatter band is very wide.
This is more difficult if the
The large scatter also tends to obscure
variables affecting fatigue properties.
Therefore, some statistical
approach to the experimental results of a cast Ti-6-4 fatigue study is
necessary for determining the importance of processing variables.
The structures selected for casting applications under study by
the Army Materials and Mechanics Research Center (AMMRC) and the Applied
Technology Laboratory (ATL) of the Army Research and Technology Labora-
- 15 tories are the Solar Titan T62T-40 gas turbine auxiliary power unit (APU)
impeller (see Fig. 1), [1-2] the UH-60A Black Hawk helicopter rotor hub
produced by Sikorsky Aircraft, [3] and a prototype centrifugal impeller
(see Fig. 2) manufactured by Detroit Diesel Allision (DDA), [4] a division of General Motors Corporation, for the Army's Technology Demonstrator
Engine (ATDE).
Solar Turbines International, currently an operating group
of International Harvester, produces the APU with a cast impeller which
would be used on the Army Black Hawk helicopter, the Air Force F-16 aircraft, and the Navy LAMPS helicopter.
Currently, a 22-pound pancake
forging is used to fashion the 2.6-pound APU impeller. [13]
The 160-
pound Black Hawk rotor hub for Sikorsky Aircraft, a division of United
Technologies Corporation, may be generated from a 250-pound casting (see
Fig. 3) instead of a 1,000-pound forging. [32]
These studies have in-
cluded castings produced by Precision Castparts Corporation (PPC),
and Titech International, Inc., with additional castings by Howmet
Turbine Components Corporation, and REM Metals Corporation, although
all four casting firms have not been involved in studies on all three
components.
The objective of this study is to determine the major variables
affecting the high cycle fatigue properties of weld repaired cast Ti-6-4
and to determine if such weld repaired material is inferior to unwelded
cast Ti-6-4 in fatigue.
The test program is statistically designed and
the results are analyzed in order to provide the most information from
the least number of samples and to correlate and confirm the results of
previous studies.
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-19
II.
PREVIOUS WORK
In many weld repair studies, a single weld pass with filler wire
is employed to fill a notch in the specimen.
Miller of TiTech [33] and
Nagon of PCC [34] used a "V"-notch normal to the bar axis.
Each conclu-
ded that welding was not detrimental to fatigue strength.
Morton and Lane [35] tested bars machined from welded plate and
found that welding was detrimental to fatigue strength.
Barice and Allen
[6] found no significant differences in fatigue of unwelded and welded
samples with spot, plug and seam welds.
Ewing and Green of DDA [10] en-
countered no adverse fatigue effects by using both a backfilled
a shallow fillerless weld.
notch and
Hammer of Solar [13] concluded that weld re-
pair was detrimental to fatigue strength based on results from blade
samples welded at the vane base and tested by reversed flexing.
Also, work by Harrigan [36] using a normal V-notch demonstrated
no change in creep mechanisms in weld repaired bars, and work by Kelly
of Howmet
37] showed no decrease in tensile strength of satisfactorily
weld repaired bars.
The study by Kelly included backfilled
normal and parallel to the axis of the bars.
notches both
He did find a decrease in
tensile strength when filler metal did not completely fill the notch.
In practice, typical weld repair procedures often require multiple
passes to build-up metal to specification.
Early passes are heat treated
as later passes are laid down, perhaps changing internal stresses and the
microstructure.
The effect of multi-pass repairs was unknown for cast
Ti-6-4 and none of the studies mentioned above reported using multi-pass
- 20 -
welds.
Because of the slow cooling rate after casting, Ti-6-4 components
have a very large PBG and colony size.
The effect of PBG and colony size
on HCF properties was discussed in the previous section.
It was not
reported whether the weld zone properties would or would not be affected
by altering the microstructure prior to weld repair by heat treatment.
It is generally accepted that HIP processing improves the HCF
properties of cast Ti-6-4 by eliminating voids. [8]
Almost all present
and future applications for dynamically stressed cast Ti-6-4 require a
HIP cycle following weld repair to ensure the integrity of the weld
metal.
No published data reported the effect of this second HIP.
As mentioned above, Kelly [37] reported that tensile strength is
not affected by weld repair if there is complete backfilling of the notch.
Other investigators found weld repair to be acceptable in terms of HCF
strength when not using either a secondary HIP [6] or a primary HIP
cycle. [33, 34]
These studies, plus a study by Taylor, Burn, and
Clarkson [38] show that porosity is not a problem in the weld zone given
proper preparation, which implies the second HIP is unnecessary.
on work by Hammer,
Based
13] there is evidence that the heat cycle during HIP
might have a deleterious effect on HCF properties, implying that the
second HIP may even be undesirable.
Repair welding induces residual stresses in Ti-6-4 which can distort the repaired structure and can have a negative effect on various
mechanical properties. [12, 13, 39]
A stress relief treatment between
900-1200°F C480-6500°C) is necessary to improve the properties. [12]
The
- 21 -
stress relief is often replaced by a higher temperature heat treatment
to alter the microstructure and further improve the properties.
Anneal-
ing for titanium alloys is carried out between 1300-1550°F (700-845°C).
Solution treating in the
(760°C) and the
+
-transus. [40]
region is performed between 1400°F
Aging follows a solution treatment or
$-treatment to relieve cooling stresses and to further alter the microstructure.
The heat treating temperature, time, and cooling rate all
affect the microstructure as mentioned previously.
At high temperatures, atmospheric purity is very important to
prevent contamination of titanium.
For this reason, heat treating is
usually accomplished in a vacuum or an inert atmosphere (usually argon).
If heat treatment is accomplished in air or if a water quench is employed,
a subsequent vacuum anneal is necessary to remove excess hydrogen while
machining or picking is necessary to remove the oxygen contaminated
c-casing.
In a water or oil quench, some components will distort before
stress relieving
or become contaminated.
A furnace or air (argon) cool
is often too slow to be an acceptable alternative.
Therefore, a rapid
argon cool is employed for these applications to maintain purity and to
rapidly cool the part without distortion.
The process involves blowing
or otherwise circulating argon through the furnace and around the heat
treating charge.
Bass [5] reported that solution treating with a water quench
produced better HCF properties than annealing with a rapid cool.
Sim-
ilarly, Hammer [13] found that solution treating with a rapid cool was
- 22 -
superior to annealing with a furnace cool.
While rapid cooling was more
desirable than slow cooling, Bartlo [41] reported HCF properties deteriorated as annealing temperature increased.
Morton and Lane [35] made
no mention of a post-weld stress relief, which may explain their conclusion that weld repair is detrimental.
Harrigan, [36] Miller, [33]
Nagan, [34] and Ewing and Green [10] all annealed specimens following weld
repair and found no decrease in HCF properties.
[9]
Silverstein and Conrad
-treated their samples and reported that rapid cooling is superior
to slow cooling.
Few studies have been published comparing the affects of various
surface finishes on HCF of cast Ti-6-4.
Bass [5] reported notch sensi-
tivity was less for annealed material than for solution treated material.
Bartlo [41] reported that heat treatment had much less effect on HCF
properties in notched bars.
Lucas [16] showed that for forged Ti-6-4,
the effect of peening on fatigue properties was only evident between
100,000 andlo million
cycles and that it had no effect on the strength
limit.
Most of the studies in the literature have used reversed bending
(R =-1.0 = maximum stress/minimum stress) methods for HCF studies, although some used stress cycles of R =-0.1.
Bass [5] compared the two
stress cycles at both room temperature and 500°F (2600 C) and reported
that fatigue limits for R =-1.0 were lower than for R = 0.1, which is to
be expected.
Based on this, the rotating beam method would seem to be
a harsher test for HCF than tension-tension.
This study will examine many of the conflicts or omissions men-
- 23 -
tioned above.
While most investigators have reported that weld repair
of cast Ti-6-4 is unharmful to HCF properties, there are published findings to the contrary.
No reports were located examining the effects of
multiple pass welding, microstructure prior to welding, and surface
finish of weld repaired material.
Despite the general acceptance of a
post-weld HIP treatment, no reports were located showing the necessity
of this secondary HIP, and some evidence exists that implies it may even
be detrimental.
well understood.
The effects of post-weld heat treatments are also not
- 24 III.
A.
EXPERIMENTAL PROCEDURE
Statistical Design
Previous studies of fatigue behavior of cast Ti-6-4 have been ac-
complished under a wide variety of material conditions and testing procedures.
Consequently, there are some disagreements between the various
investigators as to the relative importance of particular variables, including the effect of weld repair.
This variety of opinion coupled with
the limited number of test bars available for this study led to the
choice of a statistical screening experiment.
The design aided in de-
termining the major variables influencing high cycle fatigue (HCF) behavior of weld repaired cast Ti-6-4 with the fewest number of samples.
The screening design used in this study is a 12-trial PlackettBurman design. [42]
The design contains 11 factors, resembling the
degrees of freedom in a regression analysis (see Table 1).
A factor is
assigned to each variable to be studied, leaving the unassigned factors
for determination of scatter in the data.
coded "+" and "-".
"high" and "low",
Each factor has two levels,
For continuous variables, these levels correspond to
and
for discrete variables, the levels correspond
to "on" and "off" or "type one" and "type two".
For this design, each
variable can be studied using only two levels.
With the design, the condition of each trial sample is determined
from the coding of factors assigned to the variables.
The samples are
then tested for the property under study so a quantitative result can be
assigned to each trial.
The results are used to analyze the variables
- 25 I
O
+
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+
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+
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+
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a,
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- 26 and determine the scatter of the data in this way:
a.
For each factor, the sum of all the "+" trial results is subtracted from the sum of all the "-" trial results.
b.
The difference is divided by 6 (the number of "+" or "-"
values for each factor) to determine the "effect" of the
factor.
The effect is simply the difference of the mean "+"
results and the mean "-" results.
c.
The root mean square (RMS) value for the unassigned factor
effects is calculated.
d.
The appropriate "t" distribution value for the desired confidence limits and degrees of freedom is determined (see
Table 2).
The number of degrees of freedom is equal to the
number of unassigned factors.
e.
The "t" value is multiplied by the RS
value to obtain the
minimum significant factor effect.
f.
Any variable with the absolute value of its factor effect
greater than the minimum significant factor effect is considered significant within the chosen confidence limits.
g.
The absolute values of the significant variables are ranked.
The larger the absolute value, the greater the relative influence that variable has on the property studied.
While the design is able to estimate the significance of each
variable separate of the other variables, it is unable to estimate the
interactions that may exist between the variables.
The estimates may
also be correlated to groups of two-variable interactions.
When a factor
- 27 -
Table
2
"t" DISTRIBUTION
Degrees of Freedom
80
Confidence Limits (%)
90
95
99
1
3.08
6.31
2
1.89
2.92
4.30
9.93
3
1.64
2.35
3.18
5.84
4
1.53
2.13
2.78
4.60
5
1.48
2.02
2.57
4.03
6
1.44
1.94
2.45
3.71
7
1.42
1.90
2.37
3.50
8
1.40
1.86
2.31
3.36
9
1.38
1.83
2.26
3.25
10
1.37
1.82
2.23
3.17
11
1.36
1.80
2.20
3.11
12
1.36
1.78
2.18
3.06
12.71
63.66
- 28 -
effect
is found to be significant
in comparison
to the minimum
signifi-
cant factor effect, it means the factor stands out over a pool of background noise of both interactions and experimental error.
For the purposes of this study, 7 variables were chosen to be
analyzed leaving 4 unassigned factors.
The variables included micro-
structure prior to welding, welding procedure, post-weld HIP, post-weld
heat treatment
temperature and cooling rate, finishing process and
peening.
B.
Prior Microstructure - Variable 1
For the purpose of this study, separately investment cast test
rods were purchased from Howmet of Whitehall, MI.
Each rod was approx-
imately 4.5" (11 mm) long and .625" (15.9 mm) in diameter.
The rods had
been given a HIP treatment of 1650°F (900°C) for two hours at 15 ksi
(100 MPa) and chemically milled by Howmet before shipment.
The rods had
no observable surface contamination and met the specifications for cast
Ti-6-4
(see Table
3).
The rods were labelled for identification and were examined by
x-ray radiography to ensure that they were sound.
Rods were then as-
signed processing schedules according to the statistical design (see
Table
4).
To vary the microstructure, a
-solutionizing treatment followed
by a rapid argon cool and a stress relieving anneal was employed.
This
treatment was intended to refine the colony size and platelet width.
Half the specimens were
-treated and the other half were left in the
-29-
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- 30 Table 4a
PROCESS VARIABLES
Variable
"+" Condition
"-" Condition
Initial Structure
s-treat
As-received
Bead Size
Large
Small
Pressure Treatment
HIP
None
Heat Treatment
Solution Treat and Age
Anneal
Cooling Rate
Rapid
Slow
Polish
Mechanical
Chemical
Final Treatment
Peen
None
Table 4b
ROD PROCESSING SCHEDULE
Trial
Rod Numbers
Processing*
1
1,
21
s-treat/Large Bead/STA/RC/MP
2
2,
32
B-treat/Small Bead/HIP/STA/RC/CM
3
3,
13
Large Bead/HIP/STA/SC/CM
4
4, 14
$-treat/Large Bead/HIP/@/SC/CM/Peen
5
5, 15, 25, 45
B-treat/Large Bead/@/SC/MP
6
16, 36
$-treat/Small Bead/@/RC/CM/Peen
7
37, 47
Small Bead/STA/SC/MP/Peen
8
7,
Small Bead/HIP/@/RC/MP
9
9, 19
52
10
10, 20
11
11,
12
38, 48
*STA
Large Bead/STA/RC/CM/Peen
-treat/Small Bead/HIP/STA/SC/MP/Peen
Large Bead/HIP/@/RC/MP/Peen
31
Small Bead/@/SC/CM
=
Solution
=
Anneal
Treat
RC
SC
=
=
Rapid Cool
Slow Cool
MP
=
Mechanical Polish
CM
=
Chemical
ill
anmd Age
- 31 -
as-relieved condition.
The S-treatment used was 1900°F (10500 C) for two
hours with a rapid argon cool followed by an aging cycle at 1300°F (7000 C)
for two hours
and a rapid cool.
The heat treating was accomplished using a tube furnace (see
The samples were cleaned with acetone and wrapped three in a
Fig. 4).
bundle with titanium foil.
Two bundles were heat treated in each load.
The bundles were loaded into an open-ended quartz tube along with
Ti-6-4 chips.
The foil and chips were used to protect the rods from
contamination in the system.
The quartz tube was slipped into a vycor
tube with a ball joint on one end and a tapered joint on the other.
The ball joint was joined to a sealed glass rack with valves which lead
to a vent, to pressure gauges, and to a vacuum pump.
A cap with thermo-
couple leads and a valve to an argon supply was attached to the tapered
joint.
The argon supply was welding-grade quality.
hot junction was placed between the two bundles.
The thermocouple
A fiberfrax heat shield
with heat reflecting aluminum foil ends was placed in the vycor tube
near the cap.
The apparatus was then placed in a standard three-zone
tube furnace controlled by a separate thermocouple.
After the load was placed in the vycor tube and the tube sealed
from the atmosphere, the tube was pumped down and backfilled with argon
twice.
Following the second backfill, the tube was sealed and the
load brought to temperature.
During heating, argon was vented from
the tube to maintain a nearly constant pressure of one atmosphere.
Furnace cooling was accomplished by switching off the furnace.
When rapid argon cooling was required, the glass rack vent was opened and
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- 33 and argon was forced through the tube from the argon supply, while the
The cap thermo-
furnace top was lifted to increase radiation loses.
couple was used to record the cooling rates (see the Results section for
actual rates).
Prior to welding, the rods were machined to oversized blanks (see
Fig. 5).
The resulting shavings were vapor degreased with 1,1,1-
trichloroethane and used for heat treatment chips.
The slugs removed
from the rods were used to check for contamination during processing.
One slug was included with the rods during each treatment.
Each slug
was then analyzed by Luvak, Inc. of Boylston, MA for oxygen content by
vacuum fusion and hydrogen content by hot extraction.
C.
Bead Size - Variable 2
To determine the effect of repeated weld passes on HCF life of
Ti-6-4, half of the bars were welded using a single large weld bead pass
and the other half using six smaller bead passes.
Because past work had
shown that incomplete notch filling was detrimental, no notches or filler
metal were used in this study.
The weld beads were made circumferen-
tially around the reduced section normal to the bar axis.
This reduced
the risk of contamination and incomplete filling, while allowing greater
control on weld zone and bead size.
These results were compared to re-
sults obtained with unwelded specimens to find detrimental effects.
The simulated weld repair was done in MIT's welding laboratory.
The blanks were held by one end in a rotating chuck connected to a
variable speed motor.
The torch was clamped to a stand and positioned
--' OO' i
0
~.
i*
' C,
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rX
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is
4
[
0
.2i
o
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coL
r~~~~~~~~~~~~
~~-·---.X------------~~
0
0)
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;
- 35 -
The entire apparatus was
above the area on the blank to be welded.
placed in a glove box with an argon atmosphere.
The blanks were cleaned prior to welding using acetone and methylethyl-ketone.
To simulate weld repair, a single bead was run around the
reduced section of the blank using a ramped and pulsed arc program (for
The gas used to fill the glove box and used for
parameters see Table 5).
Atmosphere purity was
the shielding gas was ultra-pure (99.9%) argon.
confirmed by checking for discoloration in a spot weld made on a test
The blanks were allowed to cool in the glove box atmosphere be-
slug.
fore subsequent welding or removal from the box.
For the multi-pass
simulation, subsequent weld passes were made at lower current, but otherwise using the same welding conditions (see Table 5 and Fig. 6).
D.
Hot Isostatic Press - Variable 3
To study the effect of the secondary
samples were given a HIP treatment
HIP cycle, half the
using standard practice for Ti-6-4
The blanks designated for a HIP treatment were
following weld repair.
sent to Industrial Materials Technology Corporation (IMT) of Andover,
MA.
The blanks were given a standard 16500 F (900°C) treatment at 15 ksi
(100 MPa) for two hours.
E.
Heat Treatment - Variables 4 and 5
Because Sikorsky had reported no difficulty in
-treating and
water quenching the cast rotor hub to produce the desired microstructure
without distortion, [32] the heat treatment portion of this study concen-
- 36 -
Table
5
WELDING PARAMETERS
For all welds:
Electrode diameter
=
3/32" (1.6 mm)
Electrode type
=
Tungsten with 2% Th
Electrode tip angle
=
90 deg.
Electrode tilt angle
=
0 deg.
Tip to cup distance
=
.15"
Tip to work distance
X
.1" (2.5 mm)
Cup diameter
=
15/16"
Torch gas flow
=
20 cfh Ar (9.4 l/min)
Welding speed
=
10 rmp = 10 ipm (4.2 mm/sec)
Arc initiation
=
high-frequency
Peak voltage
=
15 volts
Peak current (first pass)
=
200 amps
Initial current
=
5%
Time at initial current
=
5.5 sec
Ramp-up time
=
1 sec
Time at peak current
=
6.2 sec
Ramp-down time
=
2
Final current
=
2%
Time at final current
=
4 sec
Background current
=
15%
Pulses
=
2 per sec
Time at pulse peaks
=
15%
=
150 amps
Peak current (fourth, fifth and sixth pass) =
100 amps
(3.8 mm)
(1.6 cm)
Ramp function:
sec
Pulse function:
For multiple pass welds:
Peak current (second and third pass)
Cooling time between passes
=
1 hr
(minimum)
L)
Cn
en
c
CY)
cU
co
C,
(L
at.
,)
a.
._to
0.
o
c'0
0.
CD
o)
ac1
0)
IL.
- 38 -
Solar [13] and DDA [43] both
trated on the more sensitive impellers.
reported difficulty with distortion on cooling from the B-treatment temperature and during water quenching from a solution treatment.
There-
fore, this study was limited to annealing and solution treating the
specimens, with the maximum cooling rate being that of a rapid argon
cool.
Variable 4 was chosen as the time and temperature variable with
variable 5 the cooling rate variable.
To study time and temperature,
To
half the bars were annealed and half were solution treated and aged.
study cooling rate, half were furnace cooled and half were rapid argon
cooled.
All blanks were given their designated heat treatment using the
tube furnace and procedure already described.
The solution treated
blanks were heat treated for 1.5 hours at 17500 F (9550 C), cooled, and
then aged at 950° F (510°C) for four hours.
heated to 1350F
F.
The annealed blanks were
(730°C) for two hours and cooled.
Surface Finish - Variables 6 and 7
Variable 6 was chosen to study the effect of a chemically milled
surface finish.
All bars were polished to a similar finish before
Thus half the samples remained in a mechanically
chemical milling.
polished condition.
Variable 7 was chosen to study peening effects.
Half of all the bars were glass bead peened after the final polish and
half were left in the polished condition.
All blanks
were machined
to final
dimensions
at AMMRC
(see Fig.
7).
0
0
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ac
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- 40 -
The specimens were then hand polished in order to achieve a fine surface
finish and to eliminate transverse scratches.
Specimens which were to
be chemically milled to final dimensions were machined slightly oversize before milling.
Chemical milling was done in a heated acid bath. [44]
The acid
consisted of 10% nitric and 3% HF in deionized water and was heated to
115°F (45°C).
Each specimen was milled separately using a small rotat-
ing tumbler to provide uniformity.
Designated bars were sent to Metals Improvement Company (MIC) of
Lynn, MA for peening.
The peening was accomplished using a slurry with
AB size glass beads, attaining an Almen intensity of .010-.012 N2 , as
per specification AMS 2430H.
G.
Testing Methods
The fatigue mode of impeller blades is considered to be closest
to a flexing motion.
13,
43]
The rotor hub has a variety of different
fatigue modes including flexing motions. [45]
The limitations of the
fatigue testing equipment provided the choice of testing in tensiontension or on a rotating beam fatigue machine.
It was desired to increase the fatigue cycle frequency as high
as possible both to better simulate the actual operating conditions, and
to decrease the testing time.
This desire plus the flexing motion in
impeller blade fatigue were the reasons for choosing R.R. Moore rotating
beam machines.
Prior to testing, the RMS surface finish of the specimens was
- 41 -
measured with a profilometer and internal defects were located by x-ray
radiography.
The specimens were each tested on one of two R.R. Moore
machines at 72.5 ksi (500 MPa) and 10,000 rpm (167 Hz).
Broken specimens were sectioned for metallography.
One half of
each specimen was sectioned logitudinally to observe the internal microstructure, while the other was sectioned transversely to observe the fracture
surface.
The microstructure was revealed by mechanically polishing and
etching in a nitric-HF solution.
The fracture and specimen surfaces were
examined in a scanning electron microscope (SEM).
Colony size measurements were made from micrographs of the base
metal in each trial specimen pair.
A line intercept method was employed
to calculate the mean intercept length (L3 ), which is a measure of the
average colony size. [46, 47]
A colony boundry was assumed to be located
wherever the prevailing platelet orientation changed in the micrograph.
For comparison to the welded specimens, ten as-received rods were
machined and mechanically polished.
Two rods were fatigue tested at
each of five stress levels: 72.5 ksi (500 MPa), 65 ksi (450 MPa), 55 ksi
(380 MPa), 50 ksi (345 MPa), and 47.5 ksi (325 MPa).
- 42 -
IV.
RESULTS
Two specimens were processed according to trial 5 of the statisti-treat, large bead, no HIP, anneal and slow cool,
cal design, namely
mechanically polished and no peen.
They were each tested at two low stress
levels, 62.5 ksi (430 MPa) and 65 ksi (450 MPA), until run-out and then run
until failure at 70 ksi (480 MPa).
Both failed at approximately 2.5
million cycles, one failing in the threads.
This procedure made these
specimens invalid for inclusion in the design, so two additional specimens
were prepared to replace them.
The processing of the additional specimens
only approximated the processing of the original specimens since much of
the welding and heat treatment equipment had to be reassembled.
The results of the R.R. Moore testing (see Table 6) could not be analyzed directly using the screening design because the replacement trial
5 specimens ran-out.
A variety of different values for trial 5 (see Appen-
dix I) did not greatly affect the results of the statistics, therefore a
value of 10 million cycles was chosen.
By calculating the statistics (see
Table 7), surface finish is well over 90% significant as a variable in HCF
life with heat treatment time and temperature nearly 80% significant.
Mechanically polishing is shown superior to a chemical milling for HCF life
and annealing is shown superior to solution treatment.
X-ray radiography showed no indications of voids either in the
as-received or post-processed material.
The as-received material showed
indications of what may have been inclusions or segregation, however the
post-processed material showed no imperfections.
RMS surface measure-
- 43 -
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CD
e0
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coJ
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--)
(n
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ul
(
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a)
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1:
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a
45 -
ments (see Table 8) showed the peened surface to be much rougher than
the chemically milled or mechanically polished surfaces.
Mechanical
polishing was only slightly finer than chemical milling.
Chemical analy-
sis of the slugs (see Table 9) showed no significant oxygen or hydrogen
contamination in any of the processes.
The time to cool from the treatment temperature to 1000°F (540°C)
was measured for each heat treatment.
was slightly over 3 minutes from 1900F
The time of rapid argon cooling
(1050°C), slightly under 2 minutes
for 1750°F (955°C), approximately 1.5 minutes from 13500 F (730°C) and
approximately 1 minute from 1300°F (700°C).
The time of furnace cooling
was under 2 hours from 1750°F (955°C) and under 1 hour from 1350°F (730°C).
Approximately 45 minutes was required to cool from the HIP temperature of
1650°F (900°C).
Microstructural analysis revealed that all failed specimens, except the specimens for trials 2 and 3, fractured outside of the weld zone.
Trials 2 and 3 were both chemically milled and had the lowest HCF lives.
The microstructure of the weld metal (see Figs. 8-11) was similar among all
the specimens.
The PBG size of the weld metal was generally very large and
the morphology was a basketweave structure with little or no GBa.
The base
metal in contrast had thicker GBa, larger colonies, and a smaller PBG size
(see Fig. 12-15).
Weld penetration was complete through the thickness of
the reduced section.
There was no visible change in microstructure across
the weld metal due to the multiple passes or in the base metal due to the
heat affected zone (HAZ), although a transition from weld metal to base
metal was observed (see Fig. 16).
- 46 -
Table
8
SPECIMEN SURFACE FINISHES
Specimen Number
RMS Range
7
5-6
6-8
4-5
15
5-6
21
5-6
4-6
1
5
52
2
3
13
32
38
48
4
9
14
16
19
36
10
11
20
31
37
47
15-19
18-22
15-23
11 -14
13-18
11-14
40-65
45-65
50-65
35-45
35-45
30-50
45-60
60-80
50-80
40-65
35-55
55-70
- 47 -
Table
9
SLUG ANALYSIS
Analysis
Treatment*
(%)
0
H
As-received
1
.21
.0048
As-received
2
.18
.0051
As-received
3
.19
.0050
S-treat (load 1)
.21
.0050
$-treat (load 2)
.21
.0074
HIP
.20
.0089
STA/RC
.19
.0052
STA/SC
.18
.0060
@/RC
.21
.0060
@/SC
.19
.0042
Chemically Milled
.19
.0054
*STA
=
Solution Treat And Age
@
=
Anneal
RC
SC
=
=
Rapid Cool
Slow Cool
i
Figure 8: Specimen 4 base metal. (optical,
HF-nitric
etch)
. mm
l i'ur e
):
Specimen
19 base metal
(optical, HF-nitric
etc }
imrrm
1..'
:·,·.
?I
··;;
·
:·T'--:-·
i
·- ic-:··-·:-.1,s
,,
5-·?
X
'".: -)
···'i.·i-":
;··;··'""··i.'-1;''"
·
nt
··
:lf;·
2:d="
c:
·
Figure 10: Specimen 36 base metal
FiLire e 11: Sp)ccimten 38 bas(
rritali
(optical,
HF-nitric
---
('_) t i<(i,
etch)
11-mll
if'-n i tric
( t(. )
Figure 12: Specimen 1 weld metal
Figure lj:
Specimen 32 weld metal
(optical, HF-nitric etch)
-- I.rmm
Fracture
through thie weld
metal is shownat bottom (opt.ical, HF-nitric (tlch)
--
$ 1mm
Figure 14: Specimen 37 weld metal
(optical,
IIF-nitric oetch)
~.-
,;
v'
-
O-7
11
:i \;;?
\\)
Figure 1: Specimen 52 weld metal
(optical,
IIF-nitric
-- -.
I ll
.~tch)
Figure 16: Specimen 1 microstructure showing transition fr(mn
weld metal to base metal Weld metal is shown ill
the upper part of the figure The bottom o!' th,
figure shows the fracture plane across the ba~se
metal (optical, }iTF-nitric etch)
b---~
-11'!1
- 53 -
Colony size measurements of the base metal for each trial (see
Table 10) were analyzed with the statistical design (see Table 11). Cooling
rate and HIP were both over 80% significant variables with slow cooling
and HIP producing a larger colony size.
The largest colony sizes were
observed in HIP processed solution-treated and slow-cooled specimens
3 and 10).
(trials
By ranking the trials by average HCF life and comparing this rank
to mean intercept length (see Table 12), a trend is revealed.
For the
mechanically polished and the peened specimens, the larger colonies with
larger intercept lengths tend to be associated with longer HCF lives.
A trend for chemically milled specimens is not observed.
The fracture surfaces were generally sufficiently intact to make
fracture origin determinations.
All fracture origins were located on or
near the surface and usually exhibited facet-like cleavage (see Figs.
There were no indications of fractures originating at voids or
17-19).
inclusions, however most of the origin sites had been somewhat smeared
during
the fatigue
testing
(see Fig. 20).
Chemical milling removed the directionality on the outer surfaces
characteristic of the mechanically polished surfaces (see Figs. 21, 22).
Peening obscured the previous polishing procedure (see Figs. 23, 24).
X-ray emission analysis in the SEM revealed that some of the glass beads
used in peening had become imbedded in the outer peened surfaces (see
Fig. 25).
By measuring the distance of the fracture surface from the minimum diameter in the reduced section, the stress at the point of failure
- 54 -
Table 10
SPECIMEN
Trial
COLONY SIZE
MeanIntercept Length (in. xl000)
1
3.6
(0.09
mm)
2
3.9
(0.10
mm)
3
6.5
(0.17
mm)
4
5.6
(0.14 mm)
5
4.7
(0.12
mm)
6
4.4
(0.11
mm)
7
3.8
(0.10
mm)
8
3.4
(0.09 mm)
9
3.4
(0.09
10
6.4
(0.16 mm)
11
5.7
(0.14
mm)
12
5.1
(0.13
mm)
mm)
55 -
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- 56 -
Table
12
COMPARISON OF FATIGUE LIFE AND COLONY SIZE
Finish*
Trial
HCF Life (log cycles)
Mean Intercept Length
(in. x1000)
MP
5
Run-out
4.7
MP
8
6.607
3.4
MP
1
6.058
3.6
PE
10
5.917
6.4
PE
4
5.899
5.6
PE
11
5.863
5.7
PE
9
5.818
3.4
PE
6
5.752
4.4
PE
7
5.682
3.8
CM
12
4.809
5.1
CM
2
4.733
3.9
CM
3
4.625
6.5
*MP = Mechanically Polished
CM
=
PE
=
Chemically Milled
Peened
F i ,u re
17:
Specimen 10 fatigue
specimen.
Fi< ire
I '
origin
showing fracture
face-
Lower part of figure is the outer surface of tlet-
S pec imen
)(:cfr[Ftl(f
sart 0
<z
LYlI 'IlT
(SEM)
9
ftt
1mm
i gue
{ o 1 'hl:I i uer
S
or iin
sh(~wing
L
nT
t heLfue
fcra
fr;1
1,ilal'c
--
1'tu r
I i
(- X
t'l
Figure
19: Specimen 13 fatigue origin in weld metal, Lwer
part of figure is the outer surface of the
specimen.
(SEM)
i., 51m
t'aLiu:ti 2:
6pecirnen 21 fatigue
s r tace Lo'wer part
surf ace
Of
rigi.n sho)wing sinInred m:'e:
(f' titulrt' is tit c. iltttP
tthe spec inetin
( SEM )
...
}....)>
,_1
ll
Fi-Lure 21: Specimen 7 outer
surface
showing mechanically
polished condition. (SEM)
---
[,' :
,, .
0. 1111
1'~~~~~~~~~~~~~~~~
1.9.1e I':ft<
ile~
s11
(.Clt
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.i--(l
t.lt'ai'l
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mi L l (I
uter
,
( tildi Jt i (-)in
sun
ur
SLI
ce
)
i)w,
2 I1h
Figure 23: Specimen 10 outer surface showing mechanically
polished and peened condition, (SEM)
--,2mln
Plg
urc }A:
5peet.lmen 4 luter
mi.led
sulrtace
.showi.ng cthlerTli
cal Iv
and p)eened c ondit ion- ( S:N )
(a)
(b)
tiuro
2W'-:
t i ,11t,(Fl
t
9 OUL
r
S1 fr.Ca
(( -
I1
Wir 1
lld
t
tI )(-
ci
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I1
\ i 11
( j~l ) -i>tfDW.;£tiej e ~*ridivN )lf c tt l lN',v
t'tci Urf ()ri. in iIld )art )1 ttLh I r t<.-tt II ' ¼t Ir
( .
i tI',:
t I t i:
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ri'~&
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t1,
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c.
showi _l
r(APT.1[)
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tl
t,,
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in ()1,<n
'"()L iIaL the dark area
'
-. 11 c ';n1 ind(1icat ionIs
int ( bv)
(S>EW1)
§f
-
- 62 -
was calculated (see Appendix II).
A stress-log cycles (S/N) graph was
then used to plot the results (see Fig. 26).
By this method, it can
be shown that the peened specimens had HCF lives between those of the
chemically milled and most of the mechanically polished specimens.
The
results of the welded specimenswere also compared on an S/N plot with
results of the unwelded specimens (see Fig. 27).
II
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- 65 -
V.
DISCUSSION
The most important result of this study is that weld repair is not
detrimental to the HCF properties of cast Ti-6-4.
Because most trials
(all but 2 and 3) failed in the base metal instead of in the weld metal,
it is valid to extend the findings of the study to cast Ti-6-4 in general
as opposed to just the weld repaired material.
While this study did not
examine the effects of weld repair on the HAZ, the colony size and S/N
data indicates that this region is also not affected adversely.
The failure of the
-treatment to have a significant effect on the
base metal microstructure was disappointing.
ure are unknown.
The reasons for this fail-
It is possible that the heat treat furnace was unable
to heat the rods above the
-transus, but this seems unlikely since both
thermocouples were reading approximately the desired temperature.
is also an unlikely possibility that the cooling rate of the
There
-treatment
was slow enough to create the same colony structure as in the as-received
material.
A more likely possibility is that the change in microstructure
was not very great and subsequent heat treatments obscured this change.
In any case, the lack of microstructural alteration would explain the
lack of any significant effect of
-treatment prior to weld repair on
fatigue properties.
Weld bead size had no significant effect on HCF properties.
Again,
there was no observable effect of bead size on microstructure, which
would explain this result.
Subsequent heat treatments probably obscured
any differences that may have existed.
The benefit of this finding is
- 66 -
that in actual practice, the welder may use whatever procedure, within
reason, that is needed to satisfactorly repair the component.
The second HIP treatment also had no significant effect on fatigue
properties according to the HCF life screening statistics.
However, the
HIP treatment did significantly increase the average colony size of the
base metal according to the colony size screening statistics.
A large
average colony size was correlated to a longer HCF life for peened and
mechanically polished specimens.
If the secondary HIP treatment does
influence HCF properties as the colony size statistics imply, the effect
was swamped by interactions with other variables, by data scatter and by
the effects of much more significant variables.
No voids were observed using both x-ray radiography and microscopy
in either the as-received or post-processed material.
The HIP treatment
was therefore unnecessary, which may help explain the HCF results.
These
results demonstrate that only one HIP treatment is normally necessary in
Ti-6-4 casings and the treatment can be made either before or after weld
repair without any adverse effects (_assumingthe weld zone to be sound).
However, any additional HIP treatment, if necessary, would not be detrimental to HCF properties.
The correlation between larger colony size and longer fatigue life
confirms the results of Yoder and Eylon [28] for the fatigue growth rate
in a specimen with a colony size much less than the specimen thickness.
The base metal fracture origins also exhibited the expected planar colony
slip facets.
29]
The correlation between slower cooling rate and larger
colony size also conforms to previous findings. [20]
- 67 The majority of specimens showed no indications of fatigue cracking in the weld metal despite a smaller colony size (a basketweave structure) and a larger PGB size than the base metal.
The basketweave morphology
and the absence of GBa in the weld metal, both conditions unfavorable for
fatigue initiation according to work by Eylon [8 ], are the most likely
reasons for this result.
Even if fatigue crack growth rate is expected
to be faster in the weld metal, crack initiation is more favorable in the
base metal.
The heat treating results of the HCF life statistical screening are
difficult to interpret.
The numerous interactions and the predominance
of the surface finish variable cloud the statistics.
It is unjustified
to use these results to say much more than that heat treating variables
have an effect on HCF properties and that for the procedures followed,
annealing was more beneficial than solution treating.
Since previous investigators concluded that solution treating with
a rapid argon cool was more beneficial than annealing [5, 6, 13], this
last finding could be a result of the interaction of the time and temperature variable with the cooling rate variable.
For example, if anneal-
ing with either cooling rate was much more beneficial than solution treating with a slow cool, and slightly less beneficial than solution treating
with a rapid cool, then the HCF life statistics would show solution treating less beneficial than annealing.
This tendency would increase if the
rapid argon cool was not rapid enough to get the full benefit of a solution treat and age procedure.
Unfortunately, this study did not generate
enough information on heat treating to either confirm or deny this
- 68 -
hypothesis.
A similar hypothesis can be proven for the peening variable, however.
The HCF life screening statistics show no significant effect for
peening, yet an S/N plot (see Fig. 26) reveals that peening is signifiantly
more beneficial for chemically milled specimens and less beneficial for
mechanically polished specimens.
These results are explained by the
effects of surface finish and surface residual stress. The fine, directional surface finish and probably slight compressive stress of the
mechanically polished specimens achieved a longer HCF life than the
rougher, chemically milled specimens which probably had no surface stress.
The peening produced a large compressive surface stress while also greatly
roughening the surface and implanting glass inclusions in the surface
layers.
The induced stress in the peened specimens was great enough to
offset the extra surface imperfections in comparison to the chemically
milled specimens,
but not in comparison to the mechanically polished
specimens.
The fracture behavior of the chemically milled specimens without
peening was quite different from the other specimens.
Four specimens frac-
tured in the weld metal and many exhibited some major surface cracking not
associated directly with the final failure.
This indicates that chemical
milling has a deleterious effect on fatigue crack initiation resistance.
Two possible causes for this behavior are grain boundary attack and hydrogen embrittlement.
The chemical milling solution used for this study had
no surfactant and was much less in total volume than called for in the
specifications.
Although the actual milling removal was within specified
- 69 -
rates, it is possible that there was some hydrogen pick-up on the surface
layers that could not be detected by hot extraction.
would be grain boundry attack by the milling solution.
A more likely cause
Microscopy of the
surface (see Fig. 21) showed the underlaying platelet structure had been
revealed.
While this sort of localized attack was mild, it may have been
enough to help initiate fatigue cracks.
- 70 -
VI.
SUMMARY AND CONCLUSIONS
Rods were processed according to a statistical screening design to
study the effects of weld repair on HCF properties of cast Ti-6-4.
Variables studied were
-treatment before welding, welding bead size, post-
weld HIP treatment, heat treatment time and temperature, heat treatment
cooling rate, surface polish and peening.
Specimens were tested by R.R.
Moore rotating beam fatigue machines and examined by metallography,
fractography,x-ray radiography, RMS surface measurements, and chemical
analysis.
The conclusions drawn from the results of this study and other
previous work are:
1.
Weld repair is not detrimental to HCF properties of cast
Ti-6-4.
2.
Surface condition is the most critical variable in HCF life of
flexing cast Ti-6-4 components.
A finer surface finish and a
higher surface compressive stress improve HCF properties.
Mechanical polishing is superior to glass bead peening, which
is superior to chemical milling.
3.
Microstructure is also a critical variable in HCF life of
cast Ti-6-4 components.
A larger colony size in the base
metal improves HCF properties by slowing fatigue crack
growth provided the colony size is small compared to the
specimen thickness.
The weld zone resists fatigue crack
initiation due to a basketweave structure and thin GBa.
4.
No secondary HIP treatment is necessary in cast Ti-6-4 if the
- 71 -
weld zone is sound.
A HIP treatment may be made either before
or after welding without detriment to HCF properties of cast
Ti-6-4.
Multiple HIP cycles, if necessary, are also not
detrimental.
5.
Multiple pass repair welds, when compared to single pass welds,
have no effect on HCF properties of cast Ti-6-4.
- 72 VII.
POSSIBILITIES FOR FUTURE WORK
Surface condition is the most critical variable in HCF properties
of flexing cast Ti-6-4 components.
Different surface finishes and resi-
dual surface compressive stresses could be compared to find an optimum
finishing procedure.
Microstructure is also a critical variable.
Different heat treat-
ments could be compared as to their effects on HCF properties.
Cor-
relations between the microstructures produced could show the relative
importance of crack initiation as compared to crack growth rate on fatigue
properties.
This study concentrated on reversed bending fatigue.
Other
studies could concentrate on different testing conditions such as varying
fatigue cycle frequency and testing temperature, or on different testing
methods, such as tension-tension.
Studies on other titanium alloys may allow broader generalizations
about
the effects
of processing
on HCF properties
to be developed.
Of
particular interest would be generalizations about effects on cast alloy
fatigue properties.
- 73 -
Appendix
I
EFFECT OF TRIAL 5 RESULTS
The replacement specimens for trial 5 ran-out during fatigue testing.
A value for the number of cycles to failure for trial 5 (n) was
necessary for the statistical analysis, however.
The statistics were
therefore calculated for each of several assumed values of "Log n" (see
Table 13).
These values were plotted against the resulting factor effects
to show that over a wide range of assumed values, the surface finish
variable is over 90% significant and the time and temperature variable is
approximately 80% significant (see Fig. 28).
For the purposes of this
study, a value of log n = 7.0 was assumed as this is nearly run-out and
the actual value has little affect on the results.
- 74 -
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0)
- 76 -
Appendix
II
FAILURE STRESS AT FRACTURE
Almost all specimens fractured in the base metal.
Although the
proper presentation of the fatigue results would show all points on an
S/N plot at the maximum stress (72.5 ksi), it is convenient to plot the
results at failure stress for comparison to unwelded results and previous
work.
To determine the failure stress, the diameter is calculated at the
fracture plane (see Fig. 29) and compared to a list of constants. [48]
These constants are multipled by the weight applied to the ends of the
specimen in the R.R. Moore machine to get the applied stress.
A constant
moment along the reduced section is assumed for the analysis. [49]
By
measuring the distance of the fracture from the minimum cross-section and
using this method, the failure stress for each specimen was calculated
(see Table
14).
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- 79 -
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