P erform an ce M etrics and

E xp erim en tal Testing of

Sean G. Leithead
Department of Mechanical
E rosio n-R esistant C om pressor
and Aerospace Engineering,
Royal M ilitary College of Canada,
Kingston, ON K7K 7B4, Canada
B lade Coatings
e-m ail: sean.ieithead@ rmc.ca
A durability test rig for erosion-resistant gas turbine engine compressor blade coatings
was designed and commissioned. Bare and coated 17-4PH steel modified NACA 6505-
William D. E. Allan1 profile blades were spun at an average speed of 10,860 rpm and exposed to garnet sand-
Department of Mechanical
laden air for 5 h a t an average sand concentration o f 2.5 g/(m3 of air) and a blade leading
and Aerospace Engineering,
edge (LE) Mach number of 0.50. The rig was designed to represent a first stage axial
Royal M ilitary College of Canada,
compressor. Two 16 pm-thick coatings were tested: Titanium nitride (TiN) and
Kingston, ON K7K 7B4, Canada
chromium-aluminum-titanium nitride (CrAlTiN), both applied using an arc physical
e-m ail: billy.allan@ rm c.ca
vapor deposition (PVD) technique. A composite scale, defined as the Leithead-Allan-
Zhao (LAZ) score, was devised to compare the durability peiformance of bare and coated
Linruo Zhao blades based on mass-loss and blade dimension changes. The bare blades’ LAZ score
Institute for Aerospace Research, was set as a benchmark of 1.00, with the TiN-coated and CrAlTiN-coated blades obtain­
National Research C ouncil of Canada, ing respective scores of 0.69 and 0.41. A lower score identified a more erosion-resistant
Ottawa, ON K 1A0R 6, Canada coating. Major locations of blade wear included: trailing edge (TE), LE, and rear suction
e-m ail: linruo.zhao@ nrc-cnrc.gc.ca surface (SS). TE thickness was reduced, the LE became blunt, and the rear SS was
scrubbed by overtip and recirculation zone vortices. The erosion effects of secondary
flows were found to be significant. Erosion damage due to reflected particles was absent
due to a low blade solidity of 0.7. The rig is best suited for durability evaluation of
erosion-resistant coatings after (AF) being proven worthy of consideration for gas tur­
bine engines through ASTM standardized testing. [DOI: 10.1115/1.4028719]

Introduction casing. The combination of the spacing between the blades and
the interaction of the rotating blades and the casing gives rise to
One common source of foreign object damage encountered by secondary flows. These flows are important because they result in
aircraft with gas turbine engines is sand ingestion, which often performance loss in addition to that caused by profile drag and
occurs when these aircraft operate in arid or desert regions. Due to off-design inlet flow [9]. If these secondary flows become laden
the greater hardness of the sand compared to steel or titanium with sand, their characteristic vortices would lead to scrubbing
compressor blades, as well as the high engine inlet velocities, ero­ erosion on the blades’ surfaces. Although the pressure losses due
sion of the compressor blades will result if exposure time and to these secondary flows are significant, for this work, the pres­
sand concentration are severe enough. ence, nature, and effects of the secondary flows were of most
In order to mitigate sand erosion damage to fan or compressor interest.
blades, various protective coatings have been developed, tested, Due to the complexity of the flows in a rotating axial compres­
manufactured, and applied to blades. Currently, test methods for sor, the shape of axial compressor blades is carefully designed in
compressor coating evaluation using realistic conditions have order to maximize performance. Any deviation from the design
been minimal. Some full-scale tests have been performed on gas airfoil profile will result in a loss in performance. Literature
turbine engines with coated blades such as Dunn et al. [1]. How­ [1,5-8] shows that the three major areas which sustain damage
ever, this is very expensive and impractical for most organizations due to engine debris or sand ingestion are the LE, blade tip, and
to carry out. To date, the majority of testing on erosion-resistant TE. Also, increased blade surface roughness has a deleterious
coatings has been performed using stationary sandblasting of effect on performance as shown by Leipold et al. [10] and
coated coupons or blades. Examples include: Immarigeon et al. Back et al. [11].
[2] (using the ASTM G76-83 test standard, a previous version of If a rig could be built to erode rotating compressor blades using
ASTM G76-07), Gorokhovsky et al. [3] and Muboyadzhyan [4]. sand under controlled conditions, then pertinent research could be
Linear cascade testing of 3-4 stationary coated blades has also conducted to determine erosion mechanisms and roles. Addition­
been conducted: Klein and Simpson [5] and Tabakoff and Mason ally, erosion-resistant coatings could be applied to the blades,
[6]. Rotating rigs were constructed by Balan and Tabakoff [7] and tested and performance-evaluated before (BE) application on full-
Ghenaiet et al. [8], but both used aluminum blades, which are not scale and proprietary gas turbine engine axial compressors.
representative of gas turbine engine compressors. The goal of this research was to design, construct, commission,
The main benefit of the rotating rig over stationary coupon and and conduct proof-of-concept testing of a rotating erosion test rig.
linear cascade sandblasting is the presence of realistic secondary This rig would use a carefully selected, operationally representa­
flows. An axial compressor stage is a three-dimensional machine tive sand flow rate to erode rotating representative steel compres­
spinning at a high rotational speed, shrouded by a nonrotating sor blades under controlled conditions. Erosion testing was
conducted on bare 17-4PH steel blades as well as on blades coated
'Corresponding author. with titanium nitride (TiN) or a National Research Council of
Contributed by the Manufacturing Materials and Metallurgy Committee of Canada (NRC) designed chromium-aluminum-titanium nitride
ASME for publication in the J ournal of E ngineering for G as T urbines and P ower.
Manuscript received July 31, 2014; final manuscript received September 23, 2014;
(CrAlTiN) blade coating. Erosion results of the blades were then
published online December 2,2014. Editor: David Wisler. compared. The thickness of the erosion-resistant coatings was

Journal of Engineering for Gas Turbines and Power MAY 2015, Vol. 137 / 052101-1
Copyright © by ASME and the Government of Canada
 determined based on the literature. This paper provides details on
the performance metrics and erosion test results.

Experimental Test Rig

Rig Components. The Royal Military College of Canada
(RMC) turbomachinery erosion rig is a new experimental test rig
capable of rotating 16 blade assemblies at speeds up to
12,000 rpm. The blade assemblies are inserted into a vibrationally
balanced 20.32 cm (8 in.) diameter AISI 4340 steel rotor that is
driven by an aircraft gas turbine compressor (GTC) gearbox and a
20 HP electric motor. The sand is injected into the rig via a
gravity-fed hopper, with a venturi generating local low pressure at
the sand inlet point to the rig. A centrifugal blower supplied the
air mass flow to the rig. An operator station was used for rig con­ Fig. 2 Airflow velocity at the blade LE (airflow from left as per
dition monitoring tasks, such as rotor vibrations, oil flow rate and convention)
temperatures, visual oil flow and sand flow confirmation, and rotor
rpm during erosion testing. A schematic of the rig is shown in
The downstream blade had to be located at least 12% of the
Fig. 1, and a full description can be found in Ref. [12],
axial chord downstream from the upstream blade in order to mini­
mize flow disturbances near the TE of the upstream blade. This
Test Blades. In order for the RMC erosion rig to have realistic was based on Benini and Toffolo’s work [13] on optimal axial dis­
applications to in-service gas turbine engines, 17-4PH (precipita­ tances between compressor rotor and stator stages. The resulting
tion hardened) stainless steel was used for the blades. The rotor, axial spacing between the upstream and downstream blades on the
dovetail inserts and blades were electric discharge machined dovetail was set at 1.04 cm (0.408 in.), 43.9% of the axial chord.
(EDM), precisely rendering the complex shapes and geometries. A maximum distance of 24% axial chord was suggested [13J;
The blade profile was NACA 6505 with rounded LE and TE, however, this was likely due to the fact that the rotor blades in a
termed V-103. The blades were 1.27 cm (0.50 in.) tall and had a gas turbine engine’s axial compressor rotate very rapidly while
chord of 2.67 cm (1.05 in.). The dovetail inserts were 2.54 cm (1.0 the stator blades are stationary. For the RMC erosion rig, the
in.) wide and 5.84 cm (2.30 in.) long. The inlet airflow angle was upstream and downstream blades both rotated, so it was decided
set such that the air encountered the blade LE at a Odeg local that increased axial distance between the blades was acceptable
angle of attack (AOA), negating any requirement for inlet guide and provided as much of a realistic single stage compressor flow
vanes. To obviate the requirement for a larger electric motor, a pattern as possible for the upstream blade.
two-blade arrangement was devised. The front (upstream) blade The EDM process was found to leave an oxide layer on the sur­
would act as the compressor blade and the test article. The rear face of the material, which is not ideal for erosion-resistant coat­
(downstream) blade would simply redirect the flow such that it ing adhesion [12]. Therefore, AF EDM, the blades were carefully
would exit the rotor parallel to the relative incidence flow. There­ polished to remove this layer. The blades were then silver-
fore, the resulting static pressure increase across the rotor was soldered to the dovetails using an oxyacetylene torch. When this
practically zero, neglecting windage and friction losses. was performed however, oxidation re-occurred on the blades and
A blade stagger angle of 27.7 deg was selected for the upstream so they were bead blasted to remove this second residual layer.
blade, resulting in an airflow angle (a) of 47.5 deg and an axial The tensile strength of the silver solder was verified experimen­
chord of 2.36 cm (0.930 in.). Figure 2 shows a summary of the re­ tally at expected operating forces by Massouh [14] using TiN-
sultant velocity triangles at the inlet to the upstream blade. The coated steel coupons that were silver-soldered together. Figure 3
blades’ midspan was used for setting the rotational speed U. Dur­ shows a bare steel blade assembly AF silver-soldering and bead
ing erosion testing, the average achieved Mach number at the blasting. The Canadian coin shown in Fig. 3 is 23.8 mm (0.937
blade LE (MaLE) was 0.50. The rotor rpm was adjusted such that in.) in diameter.
the air encountered the blade LE at a 0 deg AOA, based on the Surface finish was also an important consideration. Surface fin­
room temperature and pressure conditions during each test. ish roughness is a measure of the deviation from a perfectly flat

0 5 2 1 0 1 -2 / Vol. 137, MAY 2015 Transactions of the ASME

 Test Configuration. A rainbow test configuration was used,
which involved an alternating bare, TiN-coated and CrAlTiN-
coated blade assembly installation pattern around the rotor such
that all of the 16 blade assemblies (6 bare, 5 TiN-coated, and 5
CrAlTiN-coated) were tested under identical conditions. Figure
4(a) shows the rainbow test pattern installed in the rotor housing
with the upstream retaining ring and screws removed. Figure 4(b)
shows a top-down perspective of three test blades side by side,
and Fig. 4(c) shows the full rotor assembly with all the blades in­
stalled. In Fig. 4(c), the bearing shown in Fig. 4(a) is resting flat
on the table. The smallest shaft connects the main rotor shaft to
the gearbox.
The garnet and air were already premixed by the time the air
reached the blades, and a cone shaped structure (not shown in Fig.
4) directed the path of the gamet/air mixture into a ring shaped
annulus passage 12.7 mm (0.50 in.) wide. The cone was installed
in an aluminum housing (see Fig. 1 for exterior view) and held
Fig. 3 Blade assembly AF silver-soldering and bead blasting: centered by three streamlined attachment posts. Immediately
(a) side view and (b) front view downstream of the cone, the annulus passage became straight for
a length of 18.3 mm (0.72 in.) to allow the flow to be oriented as
close to perpendicular to the blade LE as possible. Lastly, a
surface. The typical first stage compressor blade has a surface fin­ 6.6 mm (0.26 in.) gap remained BE the garnet-laden flow arrived
ish roughness of 0.254 pm (lOp-in.) [12], and erosion-resistant at the LE of the rotating blades and the blade/housing tip-gap.
coatings are on the order of 6-30 pm (236—1181 ^r-in.) [2-4], This gap allowed for the rotor to spin freely, taking into account
Therefore, the blade surface finish had to be close to 0.254 /mi, the thickness of the 16 screw-heads and required washer balance
but could be rougher depending on the thickness of the applied weights that secured the 1.60 mm (0.063 in.) thick retaining ring
erosion-resistant coatings. As long as the coating thickness was to the rotor (shown in Fig. 4(c)). Evidence of erosion at the root
relatively greater than the surface roughness of the blade material, and along the span of each blade supports the assumption that the
the coating would smooth out most of the blade surface’s peaks garnet particles followed the same path as the airflow.
and valleys. However, maximizing the substrate material’s
smoothness was still important for coating layer smoothness and Test Duration Determination. To determine the required ero­
adhesion. sion test duration, calculations were made based on estimates of
Surface roughness tests were conducted on the blades BE and how fast the 16 /mi-thick TiN coating would erode. These were
AF erosion testing using a SJ-400 profilometer, which was capa­ based on an expected flow velocity of approx. 151 m /s at the
ble of measuring curved surfaces accurately. Three locations were blades’ LE, which is almost double the 84 m /s velocity used in
examined: the SS near the LE of the downstream blade, the SS ASTM standard tests of 16/un-thick TiN-coated specimens by
near the TE of the upstream blade, and the pressure surface (PS) Immarigeon et al. [2], Since the ASTM standard aluminum oxide
near the LE of the downstream blade. AF averaging the results, and the garnet abrasive used for the RMC erosion rig were very
the average, root mean square (RMS), max peak height, and max similar in particle size and hardness, predictions of coating wear
valley depth values for the bare blades BE erosion were: 1.47 /(in, rate at impact angles of 90 deg and 30 deg were possible. An ero­
1.86/(m, 10.75/(m, and 5.43 pm, respectively. The average and sion equation (Eq. (1)) from the work of Sundararajan and Roy
RMS values of the EDM blades were 20% smaller than previous [17] provided a method of scaling the impact erosion rate E (the
waterjet-cut blade prototypes, and the max peak was 9% smaller. ratio of the eroded material’s mass-loss in grams to the mass of
For this work, an erosion-resistant coating thickness of 16/im the erosive particles in grams) to higher velocities. The impact ve­
(630 /(-in.) was selected, since it performed best during ASTM locity (V) is in m/s, £ cons, is a constant with units of g/g(s/m )/’,
testing of Immarigeon et al. [2]. This selected coating thickness and the exponent p is 2.4 for oblique impacts and 2.55 for normal
was sufficient to overcome the substrate material surface (head-on) impacts.
roughness.
BE coating, the blade assemblies, and specially designed hold­ E = EconstVP (1)
ers were cleaned in acetone and alcohol to remove any contami­
nants. An arc PVD method was used to deposit the 16-jum-thick
TiN or CrAlTiN erosion-resistant coatings onto the blades. The
blade assemblies were mounted on a two-axis turntable during
coating. Further details are provided in Ref. [12J.

Experimental Testing Conducted

Erosive Media Selection. Garnet UT220 abrasive from Barton,
designed for use in waterjet cutting machines, was found to be
suitable based on being safe to handle, particle size and shape,
and capability for standardization. Specifically, the average parti­
cle size of Barton 220 garnet abrasives was 82 pm, which reason­
ably matched the desert sand particle size range reported in
Davison et al. [15]. This report was based on studies of aircraft
engine sand ingestion tests in the Arizona desert (conducted by
Cowherd [16]) and Afghanistan sand ground samples. The
selected garnet particles were irregular-shaped with distinct edges.
The specific gravity of the garnet was 3.9-4.1, and the Mohs Fig. 4 Rainbow blade test pattern: (a) rotor & blades in rotor
Hardness was 7.5-8.5 [12]. For comparison, the Mohs Hardness housing, ( b) blade coatings used, and (c) rotor assem bly &
of diamonds and 17-4PH steel are 10 and approx. 5, respectively. blades

Journal of Engineering for Gas Turbines and Power M A Y 2 0 1 5 , Vol. 1 3 7 / 0 5 2 1 0 1 -3

 Taking the ratio of a higher velocity V2 to a lower velocity V ,, and from one test blade to the next would be minimal. However, the sec­
isolating for E2 (erosion rate at the higher V) results in Eq. (2): ondary flow erosion effects were postulated to be better discernible.
While the ricocheting damage mode is not to be neglected, its effects
are well described in Refs. [7] and [8]: particle reflection off the PS
and subsequent impact on the rear portion of the SS, as well as parti­
cle reflection off the SS near the LE and subsequent impact on the
Using Eq. (2) for normal (90 deg) impacts, the erosion rate was PS near the blade tip. Less is currently known of the erosion damage
expected to be 4.46 times faster at 151 m /s than at 84m /s. How­ caused by secondary flows. Overall, the RMC erosion rig’s configu­
ever, since the blade passages had to be taken into account, this ration still operationally represented conventional fans and compres­
reduced the amount of sand that would impact each blade. Once sors, where erosion in the tip and hub regions was very important,
blade size, geometry and passage spacing had been considered, it and aerodynamic efficiency was not.
was expected to take 0.69 h (41 min) for the 16pm-thick TiN coat­ The flow coefficient (<j>) (the ratio of axial inlet velocity to rota­
ing to erode completely on the LE. An equivalent coating surface tional velocity) of the rig was 0.94. This was just outside the range
area on the SS or PS of the blade (assuming oblique impacts), was of 0.3-0.9 for most axial compressors [9]. As the flow coefficient
expected to completely wear away AF 9.1 h. is decreased, the compressor stage efficiency increases, except at
A test duration of 5 h was chosen in order to observe the pro­ the very low end of the spectrum [9],
gress of coating erosion on the LE and possibly 55% of coating The diffusion factor (DF) for the upstream (compressor) blade
thickness loss on the PS and SS of the blades. If, however, second­ was calculated to be 0.85 (The DF equation can be found in Refs.
ary vortex scrubbing occurred, the erosion rate on the PS and SS [9] and [12]). This was significantly greater than the range for
of the blades could be higher in some areas. most axial compressor designs, which is usually limited to less
than 0.6 to prevent significant rises in total pressure loss and the
generation of hub-comer stall on the blade’s SS [21], However,
Test Flow Conditions and Sand Concentrations. Erosion
the RMC erosion rig used a double-bladed design to have the flow
testing was conducted at an average rotation speed of 10,860 rpm
enter and exit the rotor at the same angle, in order to prevent an
at an average axial air mass flow rate of 1.11 kg/s(2.441bm/s),
overall static pressure rise and, consequently, the requirement for
resulting in an average MaLE of 0.50 and an approximate local
a more powerful electric motor. This meant that the outlet angle
0 deg AOA. Oil-flow visualization testing was undertaken prior to
(<x2) of the upstream blade was 0 deg. Therefore, this value was a
erosion testing to determine if the blades were experiencing
factor in raising the DF value. For a typical axial compressor rotor
design incidence conditions. Erosion testing was completed in 1-h
blade stage, this outlet angle would be greater than 0 deg. Another
increments, and the sand hopper was refilled AF each test incre­
contributing factor to the high DF value was that the a of 0.7 was
ment. The rig performed consistently and reliably in terms of air
much lower than for normal axial compressors. A a nearer to 1,
mass flow rate and sand flow rate. The average deviation on the
1.5, or 2 would have reduced the DF for the upstream blade to
local AOA during erosion testing was - 0 .6 deg from zero. Per­
0.69, 0.57, or 0.51, respectively.
formance parameters for each 1-h test increment are presented in
Overall, the RMC erosion rig met the requirements to operate
Ref. [12J. The average sand concentration used for hours 1—4 of
at an acceptable MaLE and (f). Brown-out sand concentrations typi­
erosion testing was 2 .5 g /(m 3 of air). For hour 5, the sand con­
cally encountered during aircraft operations in desert regions were
centration was increased to 4 .0 g /(m 3 of air), which is considered
also achieved. Internal aerodynamic characteristic comparisons
severely limited visibility or a brown-out concentration by Davi­
with previous rotating rigs such as Balan and Tabakoff [7] and
son et al. [15]. The erosion rate for both the uncoated and coated
Ghenaiet et al. [8] are presented in Ref. [12], To summarize, the
blades at a sand concentration of 4 .0 g /(m 3 of air) was deter­
RMC erosion rig was determined to be nearly as realistic as Balan
mined to be approx, twice the erosion rate as that measured at a
and Tabakoff’s [7], based on a devised realism factor [12].
sand concentration of 2 .5 g /(m 3 of air). Therefore, results were
presented for 6 equivalent hours of erosion at a sand concentration
o f2 .5 g /(m 3 of air).
Analysis Methods
Rig Internal Aerodynamics Characteristics. The design Blade Mass/Dimension Change Performance Metrics. In
M3le f°r the V103 airfoil (NACA 6505 with rounded LEs and order to obtain a macroscopic erosion quantification, the 16 blade
TEs) is 0.67 according to Hilgenfield and Pfitzner [18]. The RMC assemblies (6 bare 17-4PH steel, 5 TiN-coated, and 5 CrAlTiN-
erosion rig was not able to achieve this value due to the backpres­ coated) were weighed BE testing, AF every hour of testing and
sure present in the rig. However, an acceptable average MaLE of upon test completion. Each blade assembly was weighed using a
0.50 was reached, and impact erosion rates could be scaled up to Scientech SA 210 scale, accurate to ±0.1 mg. The coated blade
the design Mach number using Eq. (2). This would have resulted assemblies were only weighed AF the blades had been coated.
in an erosion rate 2.0 times higher at Mach 0.67 than at Mach 0.50 The upstream end of each dovetail was etched with a serial num­
for an average p of 2.475 using Eq. (2) [12]. In terms of inlet axial ber, prior to being weighed, so that each one could be tracked.
Mach number, an average of Mach 0.34 was reached in the RMC The reason for not separating the blades from the dovetail during
erosion rig. Typical axial compressor face inlet Mach numbers for weighing AF erosion testing completion was that the removal of
aircraft gas turbine engines range between 0.4 and 0.6, the highest the blade would certainly result in extra coating or solder loss,
of which is used for engines in supersonic applications [19]. How­ invalidating any measurements. This meant that the actual mass-
ever, turboprop engines routinely operate at an inlet Mach number loss of the upstream test blade on its own could not be measured
range of 0.3—0.6 [19]. Therefore, the RMC erosion rig results are precisely. However, the relative difference in mass-loss between
still applicable to aircraft with turboprop engines operating at the each of the blade assemblies could.
lower end of this Mach region. AF the mass-loss data were obtained, percent mass-loss sus­
The solidity (<r) of the installed blades (ratio of the chord to the tained by the upstream blade of each blade assembly was esti­
pitch spacing between blades) was 0.7, which is lower than previ­ mated. It was assumed that the upstream blade experienced more
ous experiments in the literature, such as a = 2.0 for Balan and erosion than the downstream blade due to increased LE exposure
Tabakoff [7] and er= 1.0 for Ghenaiet et al. [8]. It was also lower to impact erosion. Visual examination of the coated blades AF
than the average o' of 1.4 for axial compressors in aircraft gas tur­ testing led to an estimate of 70 ± 10% of the total blade assembly
bine engines [20]. The a was limited by the robust rotor design erosion applied to the upstream blade.
and dovetail widths used in the RMC erosion rig. Therefore, the Two mass-based relations were devised based on the mass-loss
low a meant that the likelihood of ricocheting of sand particles data: % erosion rate based on time (%ERT) and erosion rate based

052101-4 / Vol. 137, MAY 2015 Transactions of the ASME

on sand (ERS). %ERT is the percent compressor blade mass-loss LEs, as well as the characteristics of the flow conditions and paths
per hour of testing at a constant sand concentration. ERS is the along the blades’ surfaces. This involved painting the surfaces of
amount of compressor blade mass-loss per total mass of sand both blades of one bare steel blade assembly with an oil paraffin-
impacting one blade and passing through one blade passage graphite solution, then running the rig for 10 min at normal operat­
(which is 1/ 16th of the total area, since 16 blade assemblies were ing conditions, excluding sand, with all blade assemblies installed.
mounted in the rotor). Relative comparisons were termed RERT The dried oil was then left on the blades of that assembly when
and RERS, where the erosion of the bare 17-4PH steel blades was erosion testing was conducted, in order to provide clearer visual
defined as the base value. Therefore, the RERT and RERS were evidence of different areas of erosion.
unity (1.0) for the bare blades.
To determine blade dimension changes, the chord, LE thick­ Scanning Electron Microscope (SEM) Energy Dispersive
ness, TE thickness, and blade height were measured BE and AF X-Ray Analysis (EDAX). In addition to the optical microscope
5 h of erosion testing using a digital vernier caliper. The chord, pictures, SEM EDAX was performed on one TiN-coated blade
LE thickness, and TE thickness measurements were taken at three and one CrAlTiN-coated blade BE and AF erosion testing. The
locations: blade hub, midspan, and blade tip. LE and TE thickness EDAX used a maximum power beam of 20keV on locations on
measurements were taken no further than 2 mm from the LE or the PS and SS of each blade. As part of the analysis, a spectrum
TE (7% or 93% chord), respectively. Blade height measurements was produced, which showed the composition of elements in the
were taken at the LE, midchord, and TE. Average changes were area being exposed. Prior to erosion testing, the TiN-coated blade
then calculated for each test blade. These average changes were spectrum detected the presence of Ti, N, and some trace amounts
converted into percentages and an overall average change for the of carbon and oxygen. The CrAlTiN-coated blade spectrum
6 bare 17-4PH steel blades was obtained. The same process was detected the presence of Cr, Al, Ti, N, and some trace amounts of
conducted for the 5 TiN-coated and the 5 CrAlTiN-coated blades. carbon and oxygen prior to erosion testing. These spectra con­
Once these overall averages were calculated, they were scaled for firmed the coating compositions, but also that the electron beam
the 6 equivalent hours of erosion at a sand concentration of could not penetrate the 16/tm-thick coating, since no iron was
2.5g/(m 3 of air) to obtain percent change per hour. Table 1 detected. Therefore, AF erosion testing, if a sufficient coating
shows the resulting performance metrics and their respective thickness was eroded, the EDAX would begin to show the pres­
descriptions. For the relative comparison metrics (RECRR, ence of iron in the spectrum (since the blades were made of 17-
RTETRR, RLETIR, and RHRR), the erosion of the bare 17-4PH 4PH steel). Unfortunately, the threshold coating thickness at
steel blades was defined as the base value, therefore they were all which iron would begin to appear in the spectrum was unknown.
equal to unity for the bare blades. Making several different coating thickness samples in an attempt
Uncertainty values for equations of the form / = xy/z were cal­ to find this threshold was cost- and time-prohibitive. However, a
culated using the partial differential method [22]. Uncertainties on relative comparison could be made, based on the different spec­
averages were calculated using the standard uncertainty method, tra’s iron concentrations for different locations on a blade.
and uncertainties on sums were calculated using the summation in
quadrature method (both from Ref. [23]).
Test Results and Discussion
Visual Erosion Observation Methods. Erosion patterns were Qualitative Blade Dimension/Geometrical Changes. The
observed using both the naked eye and an optical microscope. The LEs of the bare, TiN-coated, and CrAlTiN-coated blades all
various lighting parameters used to take each optical microscope clearly sustained impact damage. Figure 5 shows photos of the LE
photo were recorded so that BE and AF photos of the same loca­ of bare, TiN-coated, and CrAlTiN-coated blades BE and AF 5 h
tion used the same lighting parameters. Photos were taken of the of erosion testing. The severity and nature of the damage differed
following parts of the upstream (test) blades: PS rear-half, PS for all three types. The LE of the bare 17-4PH steel blade became
front-half and LE, SS front-half, SS rear-half, and blade-tip pro­ bowed and more blunt (Fig. 5(b)). This indicated that the velocity
file. For the downstream blade, less magnified photos were taken gradient of the airflow conformed to a duct flow, where the veloc­
of the entire PS and SS. Optical microscope photos and visual ity is greater in the centre and decreases toward the walls. LE
descriptions using the naked eye were taken AF every hour of blunting was commonly observed in previous compressor blade
testing. Representative blade photos not shown in this paper can profile erosion research [5-8].
be found in Ref. [12]. For the TiN-coated blade, the LE became polished and slightly
Prior to erosion testing, oil-flow visualization was conducted to more blunt (Fig. 5(d)). This resulted in a greater LE thickness
confirm the intended Odeg local AOA at the upstream blades’ increase than for the bare blades, likely due to the protection

Table 1 RMC erosion rig blade erosion perform ance m etrics (PM )a

PM Units Description

%ERT % /h % blade mass-loss rate

RERT — Relative % E R T ( % E R T coated b la d e /% E R T b are blade)
ERS (g /g )/h Erosion rate based on sand (blade mass-loss per
total sand mass impacting one blade and
passing through one blade passage)
RERS — Relative E R S ( E R S coa,ed b lad e /E R S b are blade)
%ECRR % /h % effective chord reduction rate
RECRR — Relative %ECRR(%ECRRcoated blade/%ECRRbarc blade)
%TETRR % /h % TE thickness reduction rate
RTETRR — Relative % T E T R R ( % T E T R R coa,ed biad e / % T E T R R bare blade)
% /h % L E thickness increase rate
%LETIR
RLETIR — Relative % L E T I R ( % L E T I R coated b l a d e /% L E T I R b are blade)
%HRR % /h % height reduction rate
RHRR — Relative % H R R ( % H R R coated b l a d e /% H R R b are blade)

aThe erosion of the bare 17-4PH steel blades was defined as the baseline value, meaning that RERT, RERS, RECRR, RTETRR, RLETIR. and RHRR
were all unity for the bare 17-4PH steel blades.

Journal of Engineering for Gas Turbines and Power MAY 2015, Vol. 137 / 0 5 2 1 0 1 -5
 Fig. 5 Blade LE & forward PS region photos BE & AF 5 h of erosion (airflow from left to right,
scale: 1 mm per increm ent): bare 17-4PH steel (a) BE and (b) AF; TiN-coated (c) BE and Id) AF;
CrAITiN-coated (e) BE and ( f) AF (adapted from Ref. [12])

provided by the remaining TiN coating on the SS and PS just of the blade. These characteristics are even clearer in Fig. 6(c) AF
downstream from the LE. For the bare blades, at the same time 5 h of erosion testing.
that the LE was becoming more blunt, the polishing of the front Sections A-A, B-B, and C-C from Fig. 6(b) are sketched in
regions of the SS and PS eroded away some of the LE thickness. Figs. l(a)-l(c), respectively, to describe the flow characteristics
This made its overall increase in bluntness less apparent. The at each section. Section A-A shows evidence of a separation
inward bow at midspan of the TiN-coated blades (Fig. 5(d)) was
less apparent than that of the bare blades. This indicated that the
TiN-coating on the LE provided some protection for the steel sub­
strate material. The inward bow at the midspan also contributed to
the greatest chord length reduction occurring at the blade midspan
for both the bare and coated blades.
For the CrAITiN-coated blade, the LE sustained an irregular
erosion pattern, with an overall LE radius increase (Fig. 5(f)).
This resulted in almost the same LE thickness increase as that
encountered on the bare blades (according to quantitative results
outlined later in this paper). Erosion was greater than that experi­
enced by the TiN-coated blade, but less than that encountered by
the bare steel blade. There was no noticeable bow at the midspan.
This erosion pattern indicated that the CrAlTiN-coating on the LE
provided some protection for the steel substrate material, but less
than that provided by the TiN coating. Chromium and aluminum
have lower hardness characteristics than titanium; therefore, the
irregular erosion pattern could have been due to the use of three
types of metals in the coating, the distribution of which could
have led to lower adhesion capabilities in different regions of the
LE. An irregular erosion pattern would almost certainly increase
the performance losses compared to those for a uniformly eroded
blade LE. The more the LE retains its design profile, the lower the
performance losses. This is consistent with previous compressor
blade profile erosion research [7,8,24,25].
The SS of the bare, TiN-coated and CrAITiN-coated blades sus­
tained erosion damage. However, the severity and nature of the
damage differed for all three blade groups. Figures 6(a)-6(c)
show the SS of a bare 17-4PH steel blade AF oil-flow visualiza­
tion testing (no sand used), AF 1 h of erosion testing (annotated),
and AF 5 h of erosion testing, respectively. Figure 7 shows cross
section sketches of the proposed flow phenomena around the
blade at the annotated sections identified in Fig. 6(b).
In Fig. 6(b), Point 1 identifies a polished surface up to the maxi­
mum camber point (approx. 50% chord) of the blade. This shows
that the sand was likely entrained in the boundary layer and pol­
ished the surface at shallow angles. AF the midchord point, the
flow separated from the surface (Point 2), formed a separation
bubble, and reattached at Point 2', as evidenced by less residual
Fig. 6 Bare 17-4PH steel blade SS (airflow from left to right,
oil downstream of Point 2' in the blade midspan region. Point 3 scale: 1 mm per increment): (a) oil-flow visualization, ( b) AF 1 h
corresponds to a darker area caused by scrubbing erosion from of erosion (annotated), and (c) AF 5 h of erosion (adapted from
overtip vortices which travelled from the PS over the tip to the SS Ref. [12])

052101-6 / Vol. 137, MAY 2015 Transactions of the ASME

 S e p a ra tio n b u b b le higher a near 1.5, resulting in a lower DF, the recirculating flow
at the rear midspan region of the SS would not likely be present,
as found by Lei et al. [28]. Figure 8 shows the forward and rear
SS of bare and coated blades BE and AF 5 h of erosion testing.
The effects of the recirculating sand-laden flow scmbbing on
the SS and the polishing on the PS combined to reduce the blades’
TE thickness. The TE thickness reduction was more apparent on
the bare 17-4PH steel blades than for either of the two types of
coated blades. The initial tip-gap for all blades was 0.51-0.76 mm
(0.02-0.03 in.), which was 1.9-2.9% of the initial blade chord
(4_6% of initial blade height). This was controlled by using
BONDO autobody filler material in the machined rotor housing
groove. AF testing, the BONDO filler appeared to have main­
S e p a ra tio n b u b b le
tained its initial thickness. This tip-gap range was similar to that
studied by Tang et al. [29], Therefore, overtip vortices traveling
from the PS through the tip-gap to the SS were expected. These
sand-laden vortices would cause scmbbing erosion near the TE in
the tip region of the SS. The reduction in the TE thickness is evi­
dent in Fig. 9, which shows photos of the rear portion of the bare,
Fig. 7 P ro p o s e d flo w p h e n o m e n a o c c u rrin g at s e c tio n s a n n o ­ TiN-coated and CrAlTiN-coated blade tips BE and AF 5 h of
ta te d in Fig. 6(£>): (a) S e c tio n A -A , ( b) S e c tio n B -B , a n d (c) S e c ­ erosion testing.
tio n C -C (a d a p te d fro m R ef. [12])
For the TiN-coated blade, the TE thickness was substantially
reduced (Fig. 9(d)), but not to the extent as for the bare steel blade
bubble immediately downstream of Point 2, with a turbulent flow (Fig. 9(b)). Furthermore, the tip on the PS between Points 1 and 2
reattachment. Section B-B shows evidence of a separation bubble in Fig. 9(d) was rounded AF erosion, denoting the presence of
immediately downstream of Point 2, followed by a turbulent flow overtip flow erosion in this region. For the CrAlTiN-coated blade,
reattachment and a second separation point immediately prior to the TE thickness had been reduced slightly (Fig. 9(f)), but not to
Point 4. The sand-laden recirculation at Point 4 likely contributed the same degree as that of the bare steel blade nor the TiN-coated
to increased scrubbing erosion, shown by the darker color at this blade. For both the bare and coated blades, the rest of the blade
location. Section C-C shows evidence of a separation bubble thickness, except for the LE, remained effectively constant. The
immediately downstream of Point 2, which merges with the hub- full overtip surface had a more polished appearance for all three,
comer stall. This is shown by the large amount of residual oil which indicates erosion by overtip flow.
along the Section C-C line from Point 2 to Point 5 in Fig. 6(c). For the coated and uncoated blades, the height was reduced the
In compressor cascades, the transition point from laminar to most at the TE, proposed to be due to the location of the overtip
turbulent flow on the SS often occurs by means of a laminar sepa­ vortex in that region. However, it appears that the overtip vortex
ration bubble [19]. The significance of the bubble is that free shear contributed more to reduction of the TE thickness than in reduc­
layers, such as that over the bubble, are very unstable and become tion of the blade height, shown by the difference between the
turbulent at an earlier chordwise position than would have %TETRR and %HRR values in Table 2.
occurred for an attached boundary layer. Downstream of the bub­
ble, the flow in the shear layer becomes turbulent and reattaches
Quantitative Blade Erosion Results
[9,26], Blade LE geometry changes, which resulted during erosion
testing, can cause separation bubbles on the SS due to a larger Unprocessed Blade Assembly Mass-Loss. Raw mass-loss data
acceleration around the LE, followed by a localized deceleration was obtained for each blade assembly, each of which consisted of
[25]. Based upon prior linear cascade research, the transition the upstream (test) blade, downstream blade and dovetail. Figure
Reynolds number based on the chord (Rec) is typically 250,000 10 shows the unprocessed blade assembly cumulative mass-loss
[27], Therefore, a transition from a laminar to a turbulent bound­ in grams during erosion testing. Each line on the graph was gener­
ary layer likely occurred on the SS since the Rec for the blades ated from the average mass-loss values for the 6 bare blades, 5
was approx. 260,000 (with Re = 250,000 at 97% chord). However, TiN-coated blades, and 5 CrAlTiN-coated blades, respectively.
the upstream shift in the transition zone from 97% chord to Point Distinct differences are evident between the coated and uncoated
2' in Fig. 6(b), approx. 60-70% chord, is quite likely due to the blades, as well as between the two types of coated blades. The
increased flow turbulence in a rotating rig versus a linear cascade. erosion rate was approx, linear, which demonstrated the repeat­
The centrifugal blower could also have introduced additional ability of the results. The change in slope for hour 5 was due to
turbulence in the flow. the increase in sand concentration to the brown-out condition of
Since DF was 0.85 for the upstream blades, much larger than 4 .0 g /(m 3 of air) recommended in Ref. [15].
the threshold value of 0.4 for hub-comer stall [21], such a stall
likely occurred, shown by the presence of a large amount of resid­ Processed Blade Mass-Loss!Dimension Changes. Processed
ual oil at Point 5 in Fig. 6(b). Lei et al. [28] determined that for mass-loss data was based on the upstream blade only, which was
compressor blade cascades with low a, a recirculation zone devel­ estimated to have sustained 70 ± 10% of the total erosion. This esti­
oped near the TE of the SS. Evidence of such a zone is shown in mation was made based on visual observation of the blade erosion
Ref. [28] for a = 1 and a Rec of 250,000, comparable to that used patterns using both the naked eye and the optical microscope pho­
for this research (260,000). Since the RMC erosion rig had a blade tos. In order to have a reference point to the mass of the upstream
a = 0.7, it can be assumed that the recirculation zone similar to blade on its own, one spare uncoated blade was weighed and the
that shown in Ref. [21] was present. This explains the dark region surface area was calculated using the blade’s Solidworks computer-
at Point 4 in Fig. 6(b), which was likely caused by recirculating aided design file. The spare uncoated blade weighed 3.0861 g, and
flow scrubbing due to a second flow separation point (Section B-B the surface area covered by the coating was 7.419 cm2 (1.15 in.").
in Fig. 7) immediately upstream of the Point 4 darkened area. The Initial upstream blade weights for the coated blades were deter­
overtip vortices at Point 3 and the hub-comer stall at Point 5 mined by adding a calculated coating mass to the nominal bare
restricted the recirculated flow scrubbing to the midspan region of blade mass using the following information: coating thickness of
the SS near the TE (as seen in Figs. 6(b) and 6(c)). This restriction 16^m, assuming the same surface area, TiN coating density of
pattern was also present in Ref. [28], For compressors with a 5.22g/cm 3 and CrAlTiN coating density of 5.42g/cm 3. The

Journal of Engineering for Gas Turbines and Power MAY 2015, Vol. 137 / 0 5 2 1 0 1 - 7
 F ig . 8 B la d e S S p h o to s B E & A F 5 h o f e r o s io n (c h o r d w is e s e c t io n s r e m o v e d fo r c la rity , a ir ­
f lo w f r o m le ft to r ig h t, s c a le : 1 m m p e r in c r e m e n t ) : b a r e 1 7 -4 P H s te e l (a ) B E a n d ( b) A F - T iN -
c o a te d (c ) B E a n d (d ) A F ; C r A IT iN - c o a t e d (e ) B E a n d ( /) A F ( a d a p t e d f r o m R e f. [1 2 ])

CrAlTiN coating density was estimated based on its chemical hours of testing and 5 or more blades per hour for which
composition as measured using energy dispersive X-ray spectros­ mass-loss and dimension change data were obtained.
copy at NRC [12]. Based on these estimates, the initial TiN Table 2 shows that there was a significant difference in the per­
and CrAlTiN coating masses were 0.062 g and 0.064 g, respec­ cent mass-loss per hour between the bare blades and the two types
tively, which represented approx. 2.0% of the initial coated blade of coated blades. This takes into account the estimate that the
mass. upstream blade sustained 70% of the total erosion. AF 6
Table 2 shows results for the overall quantitative performance equivalent hours at a sand concentration of 2.5g/(m 3 of air), the
metrics (previously identified in Table 1) for both the uncoated TiN-coated upstream blades had lost 2.3% of their initial blade
and coated blades. Table 3 shows results for the four geometry- mass (0.072 g) on average, which was greater than the initial cal­
based metrics per blade region for both the uncoated and coated culated mass of the TiN coating. It might seem that the entire
blades. Uncertainty values were low since there were 6 equivalent coating was worn off. However, inspection of Figs. 5 and 8 shows

F ig . 9 R e a r b la d e -tip p h o to s B E & A F 5 h o f e r o s io n ( a ir f lo w fr o m le ft to r ig h t, s c a le : 1 m m p e r
in c r e m e n t): b a r e 1 7 -4 P H s te e l (a ) B E a n d (b ) A F ; T iN - c o a t e d (c ) B E a n d (of) A F ; C r A IT iN - c o a t e d
(e ) B E a n d ( /) A F (a d a p te d fr o m R e f. [1 2 ])

052101-8 / Vol. 137, MAY 2015 Transactions of the ASME

T ab le 2 U n coated and coated blade co m p ariso n by m etric (0 deg T a b le 3 U n c o a te d an d co a te d b la d e c o m p a ris o n by m e tric,
local AOA, 2 .5 g /m 3 of air sand co n centration , M aLE = 0.50)a d iv id e d by b la d e re g io n (O deg local A O A , 2 . 5 g / m 3o f air sa n d
c o n c e n tra tio n , MaLE = 0 .5 0 )a,b
Performance Units Bare TiN CrAlTiN
metricb 17-4PH steel coated coated Performance Bare 17-4PH TiN CrAlTiN
metric0 steel (%/h) coated (%/h) coated (%/h)
%ERT %/h 2.06 ± 0.30 0.38 ± 0.06 0.21 ± 0.03
RERT — 1.00 ±0.21 0.18 ±0.04 0.10 ±0.02 %CRR - tip 0.17 0.18 0.09 ± 0.02
ERS xlO "4 (g/g)/h 1.84 ±0.07 0.35 ± 0.02 0.19 ±0.01 %CRR - midspan 0.27 0.25 0.15 ±0.02
RERS — 1.00 ±0.05 0.19 ±0.01 0.10 ± 0.01 %CRR - hub 0.08 0.08 0.09
%ECRR %/h 0.17 ±0.04 0.17 ±0.04 0.11 ±0.02 %TETRR - tip 6.61 2.79 0.09
RECRR — 1.00 ±0.37 0.97 ± 0.36 0.63 ±0.19 %TETRR - midspan 6.31 0.53 0.15
%TETRR %/h 6.0 ± 0.4 1.5 ±0.6 0.8 ± 0.2 %TETRR - hub 5.18 1.08 0.09
RTETRR — 1.00 ±0.08 0.24 ± 0.09 0.14 ±0.03 %LETIR - tip 0.96 2.01 1.19
%LETIR %/h 0.9 ±0.1 1.6 ±0.2 0.8 ±0.3 %LETIR - midspan 0.77 1.52 1.00
RLETIR — 1.00 ±0.1 1.7 ±0.3 0.9 ± 0.3 %LETIR - hub 0.98 1.13 0.19
%HRR %/h 0.07 ±0.01 0.06 ± 0.01 0.05 ± 0.01 %HRR- LE 0.06 0.04 0.04
RHRR — 1.0 ±0.2 0.8 ±0.1 0.6 ±0.1 %HRR - midchord 0.06 0.06 0.04
%HRR - TE 0.10 0.07 0.06
"For sample calculations and graphical forms of the data, see Ref. [12].
bFor explanations of each metric, see Table 1. "For sample calculations and graphical forms of the data, see Ref. [12],
'’Uncertainties on the values are ±0.01%/h except where indicated.
that this was not the case. Therefore, some of the substrate steel cFor explanations of each metric, see Table 1.
must have been worn away at the LE and possibly the TE on the
PS and SS. For the CrAlTiN-coated blades AF 6 equivalent hours
at a sand concentration of 2.5g/(m 3 of air), the upstream blade whereas the TiN-coated blades sustained very similar chord length
had lost 1.2% of its blade mass (0.039 g) on average, which was reductions to that of the bare blades. This is likely due to the fact
less than the initial calculated mass of the CrAlTiN coating. that most of the TiN coating had been worn off the LE AF 2 h of
Therefore, it would seem that only a portion of the coating was erosion. If the same chord reduction failure criteria of 1.5-2%
worn off, substantiated by Figs. 5 and 8. Some of the substrate used in Ref. [5] was applied, the midspan chord of the bare blades
steel was also likely worn away at the LE and possibly near the and TiN-coated blades would have reached the failure criteria
TE on the PS and SS. within the 6 equivalent hours of erosion testing.
Based on the RERT metric, both the TiN-coated and CrAlTiN- The reduction in TE thickness was the most significant blade
coated blades, respectively, performed 82% and 90% better per dimension change. In Table 2, the bare blades’ average TE thick­
hour of sand exposure than the uncoated 17-4PH steel blades. ness reduced by 6% per hour, whereas the coated blades provided
Assuming the worst case based on the uncertainty values, the substantially lower erosion rates. The significant TE thickness
coated blades were still, respectively, 57% and 67% better. This reductions were likely due to the combination of recirculation
reconfirmed the capabilities of erosion-resistant coatings in mini­ flow scrubbing and overtip vortex scrubbing erosion on the SS,
mizing blade erosion. Furthermore, this significant improvement and polishing erosion on the PS.
was achieved with a 16 pm coating thickness, which only added AF 6 equivalent hours of testing, the tip-gap increased by
2.0% to the overall blade mass. The RERS and RERT results in 0.42% of the initial blade height (or 0.40% of the initial blade
Table 2 are very similar; however, the RERS allows for a better chord length). Tang et al. [29] showed that small percentage
comparison to previous cascade and rotating rig research such as increases in the tip-gap can strengthen overtip vortices. Table 3
Refs. [5-8]. shows that the most significant blade height reduction occurred
As shown in Table 3, the CrAlTiN-coated blades had a 44% near the TE for both the bare and coated blades. This is proposed
lower midspan chord reduction compared to the bare blades, to be due to the location of the overtip vortex in that region.

Surface Roughness Changes. Initial average surface rough­

ness (Ravg) values for the bare, TiN-coated and CrAlTiN-coated
upstream blades were: 1.17 pm, 1.34 pm, and 1.64 pm, respec­
tively. These Ravg values were less than 2 pm, which was consid­
ered the minimum threshold value for rough compressor blades in
Ref. [11]. AF erosion testing, Ravg decreased on average by 13.7%
for the bare blades, 8.3% for the TiN-coated blades and 21.7% for
the CrAlTiN-coated blades. The decrease in Ravg near the TE of
the SS was consistent with the causes of TE thickness reduction:
vortex scrubbing on the SS and low angle impingement polishing
on the PS. Scrubbing vortices also have a polishing effect, which
would lead to a decrease in roughness. The largest Ravg reduction
for both the bare and coated blades was the hub-comer stall region
(26% average reduction overall and specifically a 41% reduction
for the CrAlTiN-coated blades). The largest final Ravg value was
1.46 pm (a 14% decrease), which was measured in the recircula­
tion scrubbing zone on the SS of the CrAlTiN-coated blade.
Results from Refs. [10] and [11] showed increases in Ravg, while
in this work, the opposite was determined. It is likely that the vor­
(Note: Hr #5 was at a Sand Concentration 1,6x Higher) tices were having a polishing effect and had not scrubbed the
blade surface long enough to increase the surface roughness.
F ig . 10 C u m u la tiv e u n c o a te d a n d co a te d b la d e as s e m b ly
m a s s -lo s s d u rin g e ro s io n te s tin g (M a LE = 0.5 0, 2 5 g /( m 3 o f air)
sa n d c o n c e n tra tio n (H r N o. 1-4), 4 .0 g /( m 3 o f air) sa n d c o n c e n ­ SEM EDAX Analysis. AF erosion testing, three areas of a
tra tio n (H r No. 5)) TiN-coated and CrAlTiN-coated blade were analyzed with the

Journal of Engineering for Gas Turbines and Power MAY 2015, Vol. 137 / 052101-9
 RMC SEM using EDAX: the recirculation zone scrubbed area on
the rear section of the SS (approx. 90% chord), and locations at T a b le 4 L A Z - 1 6 s c o r e r e s u lts
approx. 25% chord on the SS and 75% chord on the PS. For the
TiN-coated blade, a maximum of 2.8% of the spectra elements Test blade LAZ-16 score and uncertainty
were iron at the first area. This appeared to show that scrubbing
Bare 17-4PH steel 1.00 ±0.09
due to recirculation vortices eroded slightly more of the coating
TiN-coated3 0.69 ± 0.08
than erosion due to polishing. Figure 11 is an EDAX spectra for a CrAlTiN-coated3 0.41 ± 0 .0 6
TiN-coated blade, with the peaks being elements detected during
the EDAX scan. The larger the peak, the higher the percentage of “Coatings were applied using arc PVD at the National Research Council of
that element as compared to the total number of atoms analyzed. Canada. Substrate blade material: 17-4PH steel.
The K or L AF each element name gives information about the
atoms’ electron arrangement, which was not used in this analysis.
For the CrAlTiN-coated blade, a maximum of 1.4% of the ana­ the test results. For the gas turbine engine industry where perform­
lyzed spectra elements were iron at the 75% chord point on the ance increases of even fractions of 1% are considered significant,
PS. This showed an opposite trend to the TiN-coated blades. these results underscore the merits of erosion-resistant coatings.
Therefore, using EDAX, comparisons between the two types of
coated blades and the erosive capabilities of vortices versus shal­ Conclusions
low angle impact polishing on each coated blade were inconclu­
sive. However, the detection of iron AF erosion testing, as shown Two gas turbine erosion-resistant coatings were tested for per­
in Fig. 11(b), supported other evidence that the coatings on these formance in a novel rig designed and built at RMC. They were
areas of the blades were partially eroded. tested in a rainbow configuration against an uncoated 17-4PH
steel blade baseline under identical operating conditions. Results
were obtained periodically over 5 h of testing using visual inspec­
tion, measurement of blade assembly mass and geometry, blade
LAZ Score. In order to provide a measurable standard for the
surface roughness measurements, and SEM EDAX.
RMC erosion rig, a LAZ standard score was defined [12], Since
Wear patterns on the test blades included: impact erosion on
the coatings tested here were 16 pm-thick, the score in this case
the LE, polishing erosion on the PS and forward half of the SS,
was termed LAZ-16. Other coating thicknesses would have a
and scrubbing erosion on the SS near the TE. Both coated blades
number corresponding to their thickness in /tm since different
eroded less than the bare 17-4PH steel blades, most noticeably at
coating thicknesses might result in different LAZ scores. This
the LE and TE. The TE thickness decreased while the LE thick­
score is a normalized relative scale between 0 and 1, with the bare
ness increased due to blunting. Of all the dimensional changes,
17-4PH steel blade performance set as 1.00. A lower score
the TE thickness reduction, on the order of 6% per hour for the
denotes better erosion resistance. The LAZ score is shown in Eq.
bare blades, was the most prominent and likely the major source
(3), which combines the following six previously defined relative
of blade mass-loss. This was almost certainly due to recirculation
performance metrics:
flow scrubbing and overtip vortex scrubbing erosion on the SS,
and polishing erosion on the PS. Major chord reduction was at
. av _ RERT + RERS + RECRR + RTETRR + RLET1R + RHRR
midspan while major height reduction was at the TE. The
6 CrAlTiN-coated blades eroded less than the TiN-coated blades in
(3) all instances. However, the CrAlTiN-coated LE eroded in an
irregular shaped pattern, whereas the TiN-coated LE became
The LAZ scores for the bare, TiN-coated and CrAlTiN-coated blunter but more uniformly smooth. Overall, secondary flows
blades are presented in Table 4. were determined to be a major factor in blade erosion by sand­
When uncertainty ranges are taken into account, the TiN-coated laden air. SEM EDAX aided in detecting a reduction in erosion-
blades performed at least 14% better than the bare blades, and the resistant coating thickness in areas where the coating was not seen
CrAlTiN-coated blades performed at least 44% better than the in photos to be fully eroded.
bare blades. When compared to the TiN-coated blades, the The CrAlTiN-coated blades’ durability was superior in terms of
CrAlTiN-coated ones performed at least 14% better. It is clear lower mass-loss and the ability to better maintain its original blade
that a distinct durability performance difference was apparent in shape, and hence, performance. This was reflected in its LAZ-16

0 5 2 1 0 1 -1 0 / Vol. 137, MAY 2015 Transactions of the ASME

score of 0.41 (a lower score denotes better performance), whereas [11] Back, S„ Hobson, G. V., Song, S„ and Millsaps, K. T„ 2012, “Effects of Reyn­
olds Number and Surface Roughness Magnitude and Location on Compressor
the TiN-coated blades’ LAZ-16 score was 0.69. The bare 17-4PH Cascade Performance,” ASME J. Turbomach., 134(5), p. 051013.
steel blades had a baseline score of 1.00. It is recommended to use [12] Leithead, S„ 2013, “A Durability Test Rig and Methodology For Erosion-
ASTM standardized testing for initial coating development, and Resistant Blade Coatings in Turbomachinery," Master’s thesis. Royal Military
then for those coatings that perform well, assess their durability in College of Canada, Kingston, ON, Canada.
[13] Benini, E„ and Toffolo, A., 2007, “Innovative Procedure to Minimize
the more turbomachinery-representative conditions of the RMC Multi-Row Compressor Blade Dynamic Loading Using Rotor-Stator
turbomachinery erosion rig. Interaction Optimization,” Proc. Inst. Mech. Eng. Part A J. Power Energy,
221(1), pp. 59-66.
[14] Massouh, R„ 2012, “A Metholodolgy and Test Rig For Durability Testing of
Acknowledgment Gas Turbine Blade Erosion Coalings,” Master’s thesis. Royal Military College
of Canada, Kingston, ON, Canada.
Financial support from the Canadian Department of National [15] Davison, C. R., Chalmers, J., and Jackson, N., 2010, “Particle Concentration
Defence’s Aerospace Engineering Research Advisory Committee Ranges for Helicopter Engine Ingestion Study With Correlations to Visibility
is acknowledged, and specifically the motivation of Major D. Lit­ and Engine Performance,” National Research Council Canada, Institute for
tle. Technical support provided by Mr. R. McKellar at NRC was Aerospace Research, Ottawa, ON, Canada, Report No. LTR-GTL-2010-0031.
[16] Cowherd, C„ 2007, “Sandblaster 2 Support of See-Through Technologies for
essential. At RMC, Dr. J. Snelgrove operated the SEM and pro­ Particulate Brownout,” Midwest Research Institute, Kansas City, MO, MRI
vided the EDAX spectra, Ms. R. Massouh began and completed Project No. 110565.1.005.
the design of several of the rig’s components, and Dr. M. LaVio- [17] Sundararajan, G., and Roy, M„ 1997, “Solid Particle Erosion Behaviour of
lette and Dr. A. Asghar provided design and analysis advice. Metallic Materials at Room and Elevated Temperatures,” Elsevier Sci. Tribol.
Int., 30(5), pp. 339-359.
RMC mechanical technologists, most particularly Mr. M. Gatien, [18] Hilgenfield, L., and Pfitzner, M., 2004, “Unsteady Boundary Layer Develop­
were indispensable to the success of this program. ment Due to Wake Passing Effects on a Highly Loaded Linear Compressor
Cascade,” ASME J. Turbomach., 126(4), pp. 493-500.
[19] Walsh, P. P„ and Fletcher, P., 2004, Gas Turbine Performance, 2nd ed„
References ASME, New York.
[1] Dunn, M., Padova, C., Moller, J., and Adams, R., 1987, “Performance Deterio­ [20] Treager, I. E., ed„ 1996, Aircraft Gas Turbine Engine Technology, 3rd ed„
ration of a Turbofan and a Turbojet Engine Upon Exposure to a Dust Environ­ McGraw-Hill, New York.
ment,” ASME J. Gas Turbines Power, 109(3), pp. 336-343. [21] Mattingly, J. D„ and von Ohain, H„ 2006, Elements o f Propulsion: Gas Tur­
[2] Immarigeon, J., Chow, D., Parmeswaran, V., Au, P., Saari, H., and Koul, A. K., bines and Rockets, American Institute of Aeronautics and Astronautics, Reston,
1997, “Erosion Testing of Coatings for Aero Engine Compressor Components,” VA.
Adv. Perform. Mater., 4(4), pp. 371-388. [22] Abernethy, R. B„ and Thompson, Jr. J. W., 1973, “Uncertainty in Gas Turbine
[3] Gorokhovsky, V., Bowman, C., Wallace, J., VanVorous, D., O’Keefe, J., Measurements,” Arnold Engineering Development Center, Air Force Systems
Champagne, V„ Pepi, M„ and Tabakoff, W„ 2009, “LAFAD Hard Ceramic and Command, Arnold Air Force Station, TN, Technical Publication No. AEDC-
Cermet Coatings for Erosion Protection of Turbomachinery Components,” TR-73-5.
ASME Paper No. GT2009-59391. [23] Bell, S„ 2013, “A Beginner’s Guide to Uncertainty of Measurement,” National
[4] Muboyadzhyan, S„ 2009, “Erosion-Resistant Coatings for Gas Turbine Com­ Physical Laboratory, Teddington, UK, accessed Jan. 15, 2013, http://
pressor Blades,” Russ. Metall. (Met.). 2009(3), pp. 183-196. www.wmo.int/pages/prog/gcos/documents/gruanmanuals/UK_NPL/mgpg 11 .pdf
[5] Klein, M., and Simpson, G„ 2004, “The Development of Innovative Methods [24] Reid, L., and Urasek, D„ 1973, “Experimental Evaluation of the Effects of a
for Erosion Testing a Russian Coating on GE T64 Gas Turbine Engine Com­ Blunt Leading Edge on the Performance of a Transonic Rotor,” ASME J. Eng.
pressor Blades,” ASME Paper No. GT2004-54336. Gas Turbines Power, 95(3), pp. 199-204.
[6] Tabakoff, W.. and Mason. R„ 2007, “Dust-Induced Deterioration of Compres­ [25] Elmstrom, M. E„ Millsaps, K. T., Hobson, G. V., and Patterson, J. S.,
sor First Stage Blades in Supersonic Cascade Erosion Wind Tunnel,” Int. J. 2011, “Impact of Nonuniform Leading Edge Coatings on the Aerodynamic
Turbo Jet Eng., 24(2), pp. 85-92. Performance of Compressor Airfoils,” ASME J. Turbomach., 133(4), p.
[7] Balan, C., and Tabakoff, W„ 1984, “Axial Flow Compressor Performance 041004.
Deterioration,” AIAA Paper No. 84-1208. [26] Lou, W„ and Hourmouziadis, J„ 2000, “Separation Bubbles Under Steady and
[8] Ghenaiet, A., Tan, S., and Elder, R„ 2004, “Experimental Investigation of Axial Periodic-Unsteady Main Flow Conditions,” ASME J. Turbomach., 122(4),
Fan Erosion and Performance Degradation,” Proc. Inst. Mech. Eng. J. Power pp. 634-643.
Energy, 218(6), pp. 437-450. [27] Aungier, R. H„ 2003, Axial-Flow Compressors, American Society of Mechani­
[9] Cumpsty, N., 1989, Compressor Aerodynamics, Longman Scientific & Techni­ cal Engineers, New York.
cal, London. [28] Lei, V., Spakovszky, Z„ and Greitzer, E„ 2008, “A Criterion for Axial Com­
[10] Leipold, R., Boese, M., and Fottner, L„ 2000, “Surface Roughness Caused by pressor Hub-Comer Stall,” ASME J. Turbomach., 130(3), p. 031006.
Precision Forging on the Flow Around a Highly Loaded Compressor Cascade,” [29] Tang, G„ Simpson, R., and Tian, Q„ 2005, “Gap Size Effect on Tip-Gap
ASME J. Turbomach., 122(3), pp. 416-425. Turbulent Flow Structure,” AIAA Paper No. 2005-4024.

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